Allowable Compressive Stress at Prestress Transfer€¦ · · 2015-09-16Allowable Compressive Stress at Prestress Transfer 5. Report Date ... increasing the allowable compressive

Technical Report Documentation Page

1. Report No. FHWA/TX-09/0-5197-4

2. Government Accession No.

3. Recipient’s Catalog No.

4. Title and Subtitle Allowable Compressive Stress at Prestress Transfer

5. Report Date December 2008

6. Performing Organization Code 7. Author(s)

Brian Schnittker and Oguzhan Bayrak 8. Performing Organization Report No.

0-5197-4

9. Performing Organization Name and Address Center for Transportation Research The University of Texas at Austin 3208 Red River, Suite 200 Austin, TX 78705-2650

10. Work Unit No. (TRAIS) 11. Contract or Grant No.

0-5197

12. Sponsoring Agency Name and Address Texas Department of Transportation Research and Technology Implementation Office P.O. Box 5080 Austin, TX 78763-5080

13. Type of Report and Period Covered Technical Report 6/1/07-8/31/08

14. Sponsoring Agency Code

15. Supplementary Notes Project performed in cooperation with the Texas Department of Transportation and the Federal Highway Administration.

16. Abstract In 2004, The Texas Department of Transportation initiated Project 5197 to investigate the feasibility of increasing the allowable compressive stress limit at prestress transfer. Initially, the live load performance of 36 specimens was evaluated by Birrcher and Bayrak (TxDOT Report 5197-1, 2007). Report 5197-4 presents the subsequent research conducted based on recommendations of Birrcher and Bayrak (2007). In this portion of TxDOT Project 5197, 45 Type-C beams and 10 4B28 box beams were tested to experimentally determine their cracking load. The Type-C beams were produced in four different fabrication plants using conventionally consolidated concrete. The 10 4B28 box beams were produced in two fabrication plants using concrete mixture designs of both self consolidating concrete as well as conventional concrete. For all specimens, measured cracking loads were compared to predicted cracking loads. The data from the 45 Type-C beams and 10 box beams were added to the 36 beams investigated by Birrcher and Bayrak (2007) to compile a comprehensive set of data from 91 specimens. An appropriate maximum compressive stress limit was determined from the ability to accurately predict the load at which cracking occurred. As the maximum compressive stress at prestress transfer was increased, a decline in cracking load prediction accuracy was observed. For the specimens subjected to high compressive stresses at release (greater than 0.65f’ci), the concrete in the pre-compressed tensile zone was subjected to the non-linear inelastic range causing microcracking to occur. This non-linear behavior (due to microcracking) was unaccounted for in prestress losses or standard design equations (P/A±Mc/I). Based on the analysis of the results, an increase of the allowable compressive stress limit at prestress transfer to 0.65f’ci is justified. Additionally, the use of self consolidating concrete with a maximum compressive stress of 0.65f’ci is not recommended.

17. Key Words Allowable release stress, compressive stress limit, prestress transfer, self consolidating concrete, microcracking

18. Distribution Statement No restrictions. This document is available to the public through the National Technical Information Service, Springfield, Virginia 22161; www.ntis.gov.

19. Security Classif. (of report) Unclassified

20. Security Classif. (of this page) Unclassified

21. No. of pages 206

22. Price

Form DOT F 1700.7 (8-72) Reproduction of completed page authorized

Allowable Compressive Stress at Prestress Transfer Brian Schnittker Oguzhan Bayrak CTR Technical Report: 0-5197-4 Report Date: December 2008 Project: 0-5197 Project Title: Allowable Compressive Stress at Prestress Transfer Sponsoring Agency: Texas Department of Transportation Performing Agency: Center for Transportation Research at The University of Texas at Austin Web Address:

http://fsel.engr.utexas.edu/

Project performed in cooperation with the Texas Department of Transportation and the Federal Highway Administration.

Center for Transportation Research The University of Texas at Austin 3208 Red River Austin, TX 78705 www.utexas.edu/research/ctr Copyright (c) 2008 Center for Transportation Research The University of Texas at Austin All rights reserved Printed in the United States of America

Disclaimers Author's Disclaimer: The contents of this report reflect the views of the authors,

who are responsible for the facts and the accuracy of the data presented herein. The contents do not necessarily reflect the official view or policies of the Federal Highway Administration or the Texas Department of Transportation (TxDOT). This report does not constitute a standard, specification, or regulation.

Patent Disclaimer: There was no invention or discovery conceived or first actually reduced to practice in the course of or under this contract, including any art, method, process, machine manufacture, design or composition of matter, or any new useful improvement thereof, or any variety of plant, which is or may be patentable under the patent laws of the United States of America or any foreign country.

Notice: The United States Government and the State of Texas do not endorse products or manufacturers. If trade or manufacturers' names appear herein, it is solely because they are considered essential to the object of this report.

Engineering Disclaimer NOT INTENDED FOR CONSTRUCTION, BIDDING, OR PERMIT PURPOSES.

O. Bayrak

Research Supervisor

Acknowledgments The funds provided by the Texas Department of Transportation (TxDOT) that

made the completion of this project possible are greatly appreciated. The support of the project director Jeff Cotham along with other members of TxDOT including Keith Ramsey, Joe Roche, Jason Tucker, Graham Bettis, Randy Cox, and Andy Naranjo were also greatly appreciated. Additionally, the prestressed/precast fabricators who donated specimens and aided with the fabrication of specimens were greatly valued.

Also greatly appreciated is the precaster photography of Gil Heldenfels.

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Table of Contents

CHAPTER 1 Introduction ................................................................................................1 1.1 Allowable Compressive Stress at Prestress Transfer ...................................1 1.2 Scope of Research ........................................................................................3 1.3 Chapter Outline ............................................................................................4

CHAPTER 2 Literature Review .......................................................................................7

2.1 Overview ......................................................................................................7 2.2 Historical Background of Allowable Compressive Stress at Release .........7

2.2.1 Allowable Stresses in Reinforced Concrete .........................................8 2.2.2 Compressive Stresses at Release in Prestressed Concrete ...................8 2.2.3 Recent Research and Discussion .........................................................9

2.2.3.1 PCI Standard Design Practice 1996, 1997, and 2003 ...................9 2.2.3.2 Russell and Pang, 1997 .................................................................9 2.2.3.3 Huo and Tadros, 1997 .................................................................10 2.2.3.4 Noppakunwijai, Tadros, Ma and Mast, 2001 ..............................10 2.2.3.5 Castro, Kreger, Bayrak, Breen, and Wood, 2004 ........................11 2.2.3.6 Hale and Russell, 2006 ................................................................11 2.2.3.7 Dolan and Krohn, 2007 ...............................................................11 2.2.3.8 Birrcher and Bayrak, 2007 ..........................................................13 2.2.3.9 Summary of Recent Research .....................................................13

2.3 Mechanical Properties and Behavior of High Strength Concrete ..............14 2.3.1 Mechanical Properties of High Strength Concrete at Early Ages......15

2.3.1.1 Khan, Cook, and Mitchell, 1995 .................................................15 2.3.2 Response to Uniaxial Loading and Quantifying Internal Damage ....16

2.3.2.1 Richart, Brandtzaeg, and Brown, 1929 .......................................17 2.3.2.2 Hsu, Slate, Sturman, and Winter, 1963 .......................................18 2.3.2.3 Ngab, Slate, and Nilson, 1981 .....................................................18 2.3.2.4 Smadi, Slate, and Nilson, 1985 and 1987 ...................................19 2.3.2.5 Delibes Liniers, 1987 ..................................................................19 2.3.2.6 Gettu, Aguando, and Oliveira, 1996 ...........................................20

2.4 Properties of Self-Consolidating Concrete ................................................21 2.4.1 D’Ambrosia, Lange, and Brinks, 2005 ..............................................23 2.4.2 Ozyildirim, and Lane, 2003, and Ozyildirim, 2005 ...........................25 2.4.3 Ozyildirim and Davis, 2007 ...............................................................27 2.4.4 Gross, Yost, and Gaynor, 2007 ..........................................................28 2.4.5 Ruiz, Staton, Do, and Hale, 2007 .......................................................32 2.4.6 Burgueño and Bendert, 2007 .............................................................35 2.4.7 Naito, Parent, and Brunn, 2006 ..........................................................39

2.5 Methods to Analyze Prestress Loss ...........................................................41 2.5.1 NCHRP Report 496 Detailed Prestress Loss Method........................42

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2.5.2 AASHTO-LRFD Refined Loss of Prestress Estimate – Interim Edition 2008 ......................................................................................46

2.6 Summary ....................................................................................................48

CHAPTER 3 Test Specimens ..........................................................................................51 3.1 Overview ....................................................................................................51 3.2 Design and Fabrication of TxDOT Type-C Beams ...................................51

3.2.1 Design of Series 1 Beams ..................................................................52 3.2.2 Design of Series 2 Beams ..................................................................56 3.2.3 Fabrication of Type-C Girders ...........................................................57

3.2.3.1 Production of Girders: Fabricator A ...........................................58 3.2.3.2 Production of Girders: Fabricator B ...........................................60 3.2.3.3 Production of Girders: Fabricator C ...........................................63 3.2.3.4 Production of Girders: Fabricator D ...........................................64

3.2.4 Shipment and Storage of Type C Girders ..........................................66 3.3 Design and Fabrication of TxDOT Box Beams .........................................66

3.3.1 Design of TxDOT Box Beams ...........................................................69 3.3.2 Production of Box Beams: Fabricator E ...........................................71 3.3.3 Shipment and Storage of Box Beams ................................................74

3.4 Summary ....................................................................................................75

CHAPTER 4 Test Setup ..................................................................................................77 4.1 Overview ....................................................................................................77 4.2 Testing of Full-Scale Type-C Girders........................................................77

4.2.1 Load Protocol .....................................................................................78 4.2.2 Test Setup...........................................................................................79 4.2.3 Instrumentation and Data Acquisition ...............................................82

4.3 Testing of Full-Scale 4B28 Box Beams .....................................................84 4.3.1 Load Protocol .....................................................................................85 4.3.2 Test Setup...........................................................................................85 4.3.3 Instrumentation and Data Acquisition ...............................................89

4.4 Summary ....................................................................................................90

CHAPTER 5 Analysis of Experimental Results ...........................................................91 5.1 Overview ....................................................................................................91 5.2 Results of Type-C Beam Tests ..................................................................91

5.2.1 Measured Cracking Loads .................................................................92 5.2.2 Predicted Cracking Loads ..................................................................97

5.2.2.1 NCHRP Report 496 Method .......................................................99 5.2.2.2 AASHTO-LRFD 2008 Method .................................................104

5.3 Results of Box Beam Tests ......................................................................107 5.3.1 Measured Cracking Loads ...............................................................108 5.3.2 Predicted Cracking Loads ................................................................111

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5.3.2.1 NCHRP Report 496 Method .....................................................112 5.3.2.2 AASHTO-LRFD 2008 Method .................................................115

5.4 Observations on Self Consolidating Concrete .........................................117 5.4.1 Fabrication of Box Beam Girders ....................................................117 5.4.2 Top Flange Cracking at Release ......................................................120 5.4.3 Modulus of Elasiticty .......................................................................122 5.4.4 Camber .............................................................................................123 5.4.5 Cracking Load ..................................................................................126 5.4.6 Summary of Observations on Self Consolidating Concrete ............129

5.5 Summary ..................................................................................................129

CHAPTER 6 Summary, Conclusions, and Recommendations .................................131 6.1 Summary of Experimental Program ........................................................131 6.2 Conclusions and Recommendations ........................................................132 6.3 Recommendations for Future Work .........................................................133

APPENDIX A .................................................................................................................135

APPENDIX B .................................................................................................................155

APPENDIX C .................................................................................................................177

APPENDIX D .................................................................................................................179

BIBLIOGRAPHY ..........................................................................................................185

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List of Tables Table 2-1 Development of Allowable Stresses in Reinforced Concrete .............................8 Table 2-2 Summary of Recent Work Regarding Compressive Stress at Release ..............14 Table 2-3 Concrete mixture proportions (D’Ambrosia, 2005) ..........................................23 Table 2-4 Hardened Concrete Tests and Specifications ....................................................25 Table 2-5 Hardened Properties of Instrumented Bridge Beams (Ozyildirim and Davis,

2007) .............................................................................................................28 Table 2-6 Concrete Mixture Proportions per cubic yard (Gross et al., 2007) ...................29 Table 2-7 Concrete Mixture Designs (Ruiz et al., 2007) ...................................................32 Table 2-8 Concrete Mixture Proportions (Burgueño and Bendert, 2007) .........................35 Table 2-9 NCHRP Report 496 Material Properties Equations (Tadros et al., 2003).........43 Table 2-10 NCHRP Report 496 Equations for Prestress Loss (Tadros et al., 2003) .........45 Table 2-11 AASHTO-LRFD Equations for Material Properties (AASHTO, 2008) .........46 Table 2-12 AASHTO-LRFD Equations for Prestress Loss (AASHTO 2008) ..................48 Table 3-1 Type-C Beam Fabrication Details .....................................................................52 Table 3-2 Section Properties of a TxDOT Type-C Beam ..................................................52 Table 3-3 Calculated maximum compressive stresses at prestress transfer: Series 1

Beams............................................................................................................55 Table 3-4 Calculated maximum compressive stresses at prestress transfer: Series 2

Beams............................................................................................................57 Table 3-5 Concrete Mixture Design used by Fabricator A (per cubic yard) .....................58 Table 3-6 Beam Production Details: Fabricator A ...........................................................60 Table 3-7 Concrete Mixture Design used by Fabricator B (per cubic yard) .....................61 Table 3-8 Beam Production Details: Fabricator B ............................................................62 Table 3-9 Concrete Mixture Design used by Fabricator C (per cubic yard) .....................63 Table 3-10 Beam Production Details: Fabricator C ..........................................................64 Table 3-11 Concrete Mixture Design used by Fabricator D (per cubic yard) ...................65 Table 3-12 Beam Production Details: Fabricator D .........................................................65 Table 3-13 Box Beam Design Details ...............................................................................69 Table 3-14 Section Properties of a TxDOT 4B28 Box Beam ...........................................69 Table 3-15 Concrete Mix Designs used by Fabricator E (per cubic yard) ........................72 Table 3-16 Beam Production Details: Fabricator E ..........................................................74 Table 5-1 Measured Cracking Loads for Series 1 Type-C Beams ....................................96 Table 5-2 Measured Cracking Loads for Series 2 Type-C Beams ....................................97 Table 5-3 Calculated K1 values ........................................................................................100 Table 5-4 Box Beam Design Details ...............................................................................107 Table 5-5 Measured Cracking Loads for 4B28 Box Beam Specimens ...........................110 Table 5-6 Calculated K1 values for Box Beam test specimens ........................................113

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List of Figures Figure 1-1 TxDOT Project 5197 ..........................................................................................2 Figure 2-1 Compressive Stress at Transfer used by PCI survey respondents (adopted

from Dolan and Krohn, 2007) ......................................................................12 Figure 2-2 Theoretical stress vs. strain response for various concrete strengths ...............15 Figure 2-3 Stress-Strain behavior for a 10,000-psi concrete at early ages (Khan et al.,

1995 adopted from Birrcher and Bayrak, 2007) ...........................................16 Figure 2-4 Depiction of three stages discussed by Richart et al. 1929, adopted from

Birrcher, 2007 ...............................................................................................17 Figure 2-5 Tensile strength loss as a function of compressive stress for various

compression times and general curing conditions (Delibes Liniers, 1987). ............................................................................................................20

Figure 2-6 Typical volume fraction for different types of concrete (adopted from Bonen, 2005). ................................................................................................22

Figure 2-7 Total Shrinkage of 3 x 3 x 24.5-inch Specimens (D’Ambrosia et al., 2005) ...24 Figure 2-8 Slump Flow Test (TxDOT Research Study 5197) ...........................................26 Figure 2-9 Schematic Diagram of a U-Tube (U-Box) Test (Khayat et al., 2004) .............26 Figure 2-10 Beam Cross section and strain gauge locations (adopted from Gross et

al., 2007) .......................................................................................................29 Figure 2-11 Measured prestress loss versus time for SCC beams and HSC beams,

respectively (Gross et al., 2007) ...................................................................31 Figure 2-12 Beam Cross Section Details (adopted from Ruiz et al., 2007) .......................33 Figure 2-13 Measured prestress loss versus beam age (Ruiz et al., 2007) ........................34 Figure 2-14 Box Beam Cross Section (Burgueño and Bendert, 2007) ..............................36 Figure 2-15 Full Scale Beam Test Setups (Burgueño and Bendert, 2007) ........................37 Figure 2-16 Plan view with beam layout of M-50/US-127 bridge (Burgueño and

Bendert, 2007) ..............................................................................................38 Figure 2-17 Strain monitoring at beam bottom flange at mid-span section (Burgueño

and Bendert, 2007) ........................................................................................39 Figure 2-18 Bulb-tee beam cross section (Naito et al., 2006) ...........................................40 Figure 2-19 Stress in strands vs. time for a Prestressed Girder (Tadros et al., 2003) ........42 Figure 3-1 Strand Pattern: Series 1 Specimens .................................................................53 Figure 3-2 Type-C Beam Harped Strands .........................................................................53 Figure 3-3 Typical Harping for Fabricator C .....................................................................54 Figure 3-4 Typical Harping for Fabricator D (photographs courtesy of David

Birrcher) ........................................................................................................54 Figure 3-5 Schematic Drawing of Type-C Beam ..............................................................55 Figure 3-6 Strand Pattern: Series 2 Beams .......................................................................56 Figure 3-7 Live end of prestressing line (Fabricator A) ....................................................59 Figure 3-8 Altered Shear Reinforcement ...........................................................................61 Figure 3-9 Lifting a Type C Girder at FSEL .....................................................................66 Figure 3-10 Box Beam Void Geometry .............................................................................68

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Figure 3-11 Box Beam Centerline Strand Pattern .............................................................70 Figure 3-12 Debonding pattern for Box Beams (Not to scale) ..........................................71 Figure 3-13 Debonded strands prior to concrete placement ..............................................71 Figure 3-14 Lifting a Box Beam at FSEL ..........................................................................75 Figure 4-1 Series 1 Beam Loading Protocol ......................................................................78 Figure 4-2 Series 2 Beam Loading Protocol ......................................................................79 Figure 4-3 Type-C Beam Test Setup (Not to Scale) ..........................................................80 Figure 4-4 Picture of a Type-C Beam Flexural Test Setup ...............................................80 Figure 4-5 Midspan Region of a Type-C Beam Test .........................................................81 Figure 4-6 Type C-Beam Roller Support and Pin Support, respectively ..........................82 Figure 4-7 Linear Potentiometers Located at Midspan ......................................................83 Figure 4-8 Linear Potentiometer Located at a Support (Type-C Beam) ...........................83 Figure 4-9 Crack Documentation using a Crack Comparator Card ...................................84 Figure 4-10 Box Beam Loading Protocol ..........................................................................85 Figure 4-11 Box Beam Loading Apparatus .......................................................................86 Figure 4-12 Box Beam Roller Support and Pin Support, respectively ..............................87 Figure 4-13 Pinned connection with 10-inch long top plate ..............................................87 Figure 4-14 Box Beam Test Setup (Not to Scale) .............................................................88 Figure 4-15 Picture of Box Beam Flexural Test Setup ......................................................88 Figure 4-16 Linear Potentiometers Located at Midspan ....................................................89 Figure 4-17 Linear Potentiometer Located at a Support (Box Beam) ...............................89 Figure 5-1 Applied Load versus Midspan Deflection for Specimen CA-60-3 ..................93 Figure 5-2 Applied Load versus Midspan Deflection for Specimen CC-70-2 ..................93 Figure 5-3 Visually observed cracking load for Series 1 Beam CA-60-1 .........................94 Figure 5-4 Visually observed cracking load for Series 2 Beam CD-65-5 .........................95 Figure 5-5 Typical crack pattern at maximum applied load for Type-C Beam tests .........95 Figure 5-6 Type-C Beam Specimens: Applied Load, Shear Force, and Bending

Moment Diagrams ........................................................................................98 Figure 5-7 Estimated initial slope of load deflection plot to calculate the apparent

modulus of elasticity for CB-70-1 ..............................................................100 Figure 5-8 Comparison of measured and predicted cracking loads using NCHRP 496

losses ...........................................................................................................103 Figure 5-9 Comparison of measured and predicted cracking loads using AASHTO-

LRFD 2008 losses .......................................................................................106 Figure 5-10 Applied Load versus Midspan Deflection for Specimen BB-03

(Conventional Concrete) .............................................................................108 Figure 5-11 Applied Load versus Midspan Deflection for Specimen BB-10 (SCC) ......109 Figure 5-12 Visually observed cracking load for specimen BB-10 .................................110 Figure 5-13 4B28 Box Beams: Applied Load, Shear Force, and Bending Moment

Diagrams .....................................................................................................111 Figure 5-14 Estimated initial slope of load deflection plot to calculate the apparent

modulus of elasticity for BB-05..................................................................113

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Figure 5-15 Comparison of measured and predicted cracking loads using NCHRP 496 losses ...........................................................................................................114

Figure 5-16 Comparison of measured and predicted cracking loads using AASHTO-LRFD 2008 losses .......................................................................................116

Figure 5-17 Sample of honeycombing in box beam BB-06 ............................................118 Figure 5-18 Sample of honeycombing in box beam BB-07 ............................................118 Figure 5-19 Repaired Honeycombing of box beam BB-06 .............................................119 Figure 5-20 Repaired Honeycombing of box beam BB-07 .............................................120 Figure 5-21 Typical flexural tension cracking in the top flange of test specimens .........121 Figure 5-22 Typical flexural tension cracking in the top flange of test specimens .........121 Figure 5-23 Load versus midspan deflection for box beams fabricated with concrete

containing hard river gravel course aggregate ............................................122 Figure 5-24 Load versus midspan deflection for box beams fabricated with concrete

containing limestone coarse aggregate .......................................................123 Figure 5-25 Camber measurement at midspan for box beam BB-05 (West face) ...........124 Figure 5-26 Measured Camber in box beams fabricated with concrete containing hard

river gravel coarse aggregate ......................................................................125 Figure 5-27 Measured camber in box beams fabricated with concrete containing

limestone coarse aggregate .........................................................................125 Figure 5-28 Visually observed cracks of beam BB-04 and BB-08 respectively at an

applied load of 200-kips .............................................................................127 Figure 5-29 Visually observed cracks of beam BB-01 and BB-06 respectively at an

applied load of 180-kips .............................................................................128 Figure A-1 Sample Shop Drawing for Series 1 Type-C Beam .......................................136 Figure A-2 Sample Shop Drawing for Series 2 Type-C Beam ........................................137 Figure A-3 Sample Stress Calculations at Prestress Release for beam CA-70-1 (Series

1) .................................................................................................................139 Figure A-4 Sample Stress Calculations at Prestress Release for beam CC-70-1 (Series

2) .................................................................................................................140 Figure A-5 Prestress Loss/Cracking Load Calculations according to the NCHRP 496

procedure for Specimen CC-70-2, page 1 of 3 ...........................................141 Figure A-6 Prestress Loss/Cracking Load Calculations according to the NCHRP 496

procedure for Specimen CC-70-2, page 2 of 3 ...........................................143 Figure A-7 Prestress Loss/Cracking Load Calculations according to the NCHRP 496

procedure for Specimen CC-70-2, page 3 of 3 ...........................................143 Figure A-8 Prestress Loss/Cracking Load Calculations according to the AASHTO-

LRFD Interim 2008 procedure for Specimen CC-70-2, page 1 of 3 ..........144 Figure A-9 Prestress Loss/Cracking Load Calculations according to the AASHTO-

LRFD Interim 2008 procedure for Specimen CC-70-2, page 2 of 3 ..........145 Figure A-10 Prestress Loss/Cracking Load Calculations according to the AASHTO-

LRFD Interim 2008 procedure for Specimen CC-70-2, page 3 of 3 ..........146 Figure A-11 Sample Box Beam Shop Drawing for specimen BB-01 .............................147 Figure A-12 Sample Stress Calculations at Prestress Release for beam BB-01 ..............148

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Figure A-13 Prestress Loss/Cracking Load Calculations according to the NCHRP 496 procedure for Specimen BB-01, page 1 of 3 ..............................................149



Figure A-16 Prestress Loss/Cracking Load Calculations according to the AASHTO-LRFD Interim 2008 procedure for Specimen BB-01, page 1 of 3 .............152



Figure C-1 Static Flexural Test Information ....................................................................178 Figure D-1 Bursting crack map key for box beams .........................................................179

1

CHAPTER 1 Introduction

1.1 ALLOWABLE COMPRESSIVE STRESS AT PRESTRESS TRANSFER The maximum allowable compressive stress limit at prestress transfer was first

introduced by the American Association of State Highway and Transportation Officials (AASHTO) in the 1961 Bridge Design Specifications. Shortly thereafter, the American Concrete Institute (ACI) adopted a maximum allowable compressive stress limit at prestress transfer in the 1963 Building Design Specifications (ACI 318-63). This maximum allowable compressive stress at prestress transfer was presented in both documents as a fraction of the compressive strength of concrete. The limits prescribed by AASHTO in 1961 have not changed since their inception. In the current AASHTO-LRFD (Interim 2008) Bridge Design Specifications, the requirement concerning the allowable release stress is as follows:

The compressive stress limit for pretensioned and post-tensioned concrete components, including segmentally constructed bridges, shall be 0.60f'ci (AASHTO LRFD 2008). In the ACI 318 Building Code Requirements, the original allowable maximum

compressive stress at prestress transfer was adopted in 1963 as 0.60f’ci. This limit remained in ACI 318 until 2008 when the provision was modified to address the allowable compressive stress at the ends of simply supported members. The current code provision regarding this limit in ACI 318-08 is as follows:

Stresses in concrete immediately after prestress transfer (before time-dependent prestress losses): (a) Extreme fiber stress in compression except as permitted in (b) shall not exceed 0.60f'ci. (b) Extreme fiber stress in compression at ends of simply supported members shall not exceed 0.70f'ci… (ACI 318, 2008). In the past several years, there has been significant interest in increasing the

allowable limit of 0.60f’ci. Relaxing the maximum allowable release stress limit has many potential benefits to the fabrication and design of prestressed girders. Some of these benefits include:

• A decrease in the cycle time of precast facilities • A reduction in the number of harped strands in a given beam • A decrease in the number of debonded strands (leading to

improved durability and shear performance) • A longer span length due to an increased number of prestressing

strands in a given cross-section • The removal of “unnecessary” conservatism in current practice

Because of these potential benefits, many research projects have been conducted to determine the effects of increasing the allowable limit. One of these research projects

2

was funded by the Texas Department of Transportation (TxDOT) and conducted at the Ferguson Structural Engineering Laboratory at University of Texas at Austin. This project, denoted as Project 5197, was funded to investigate the effects of increasing the allowable stress limit in prestressed girders. TxDOT project 5197 has been ongoing since 2004 and has been conducted in two phases. Figure 1-1 displays a visual representation of the 2 phases of TxDOT Project 5197.

Figure 1-1 TxDOT Project 5197

Phase 1 of TxDOT Project 5197 consisted of two parts: • The live load evaluation of prestressed concrete beams subjected to

compressive stresses at release beyond the allowable limit of 0.60f’ci (Birrcher and Bayrak, 2007).

• The investigation of tensile cracking in short AASHTO Type IV girders (Tuchschererer, Mraz, and Bayrak, 2007).

Birrcher and Bayrak experimentally evaluated thirty-six beams of which twenty-four were scaled rectangular, tee, and inverted tee beams and the other twelve were full-scale TxDOT Type-A specimens (28-inch deep I beams). All thirty-six specimens were loaded statically to determine their cracking load. These experimentally measured cracking loads were compared to analytically predicted cracking loads using established methods to calculate prestress losses. Based on the experimentally measured cracking loads, the accuracy of all the predicted cracking loads were determined. For each specimen, the accuracy of the cracking load was plotted against the maximum

TxDOT Project 5197

Phase 1

Live Load Performance of Prestressed Girders

Report 5197-1

Birrcher and Bayrak, 2007

Tensile Cracking in Short AASHTO Type IV Girders

Report 5197-2

Tuchscherer, Mraz and Bayrak, 2007

Phase 2

Part 2:Shear Testing of Type-C Beams

Report 5197-3

Heckmann and Bayrak, 2008

Part 1:Live Load Performance of

Type-C Beams and 4B28 Box Beams

This Report

3

compressive stress at prestress transfer. Upon analysis of the data, Birrcher and Bayrak observed an unconservative decrease in cracking load prediction accuracy in specimens subjected to maximum compressive stresses (at prestress transfer) of 0.65f’ci or higher. Before endorsing a new allowable compressive stress limit at prestress transfer, Birrcher and Bayrak recommended additional testing on different section types and concrete mixture designs. This recommendation is the basis for the work presented in this report.

In addition to the static testing, four of the scaled specimens were tested under fatigue loads and an initial camber database was compiled with information from 223 pretensioned girders. The work completed by Birrcher and Bayrak (2007) is summarized in TxDOT report 5197-1.

Tuchscherer, Mraz, and Bayrak (2007) investigated the tension cracking observed at prestress transfer in short AASHTO Type IV girders. Tuchscherer et al. cast and tested several short AASHTO Type IV girders in an effort to determine an appropriate tensile stress limit in order to eliminate the observed cracking in short Type IV girders. The research by Tuchscherer et al. does not directly pertain to the research presented in this report, but is available in TxDOT report 5197-2.

Phase 2 of TxDOT Project 5197 was also completed in two parts: • Evaluation of the shear performance of Type-C beams stressed beyond the

allowable limit (Heckmann and Bayrak, 2008). • The live load performance of Type-C beams and 4B28 box beams (This

Report). Heckmann and Bayrak investigated the impact of increasing the allowable

compressive stress limit on the shear capacity of 18 full scale Type-C beam specimens (40-inch deep I beams). In essence, the project evaluated the feasibility of allowing 0.70f’ci in the end regions of prestressed Type-C girders. The information from part 2 of phase 2 of TxDOT project 5197 is available in TxDOT report 5197-3.

The live load performance of Type-C girders and 4B28 box beams (28-inch deep by 48-inch wide box beams) is the focus of this report and discussed further in section 1.2.

1.2 SCOPE OF RESEARCH In part 1 of phase 2 of TxDOT Project 5197, experimental research was

conducted based on the recommendation from Birrcher and Bayrak (2007). A thorough literature review was conducted along with experimental testing on fifty five full scale beams stressed beyond the allowable limit of 0.60f’ci. In the literature review, the historical background of the allowable release stress was presented along with recent research regarding the increase of the allowable compressive stress limit. In addition, the early age properties of high strength concrete, the properties of self consolidating concrete (SCC), and the two procedures used to estimate prestress loss in the test specimens were discussed.

The experimental program consisted of the static testing of forty-five Type-C beams and the proof testing of ten 4B28 box beams in flexure. In order to achieve a

4

range of representative data, four different fabricators produced the forty five Type-C beams. The use of several fabricators incorporated different materials and concrete mixture designs from around the state of Texas. Upon completion of the forty five Type-C flexural tests, ten 4B28 box beams were fabricated with a target maximum compressive stress at release of 0.66f’ci in order to perform proof testing on a compressive stress limit that seemed adequate for the Type-C beams. Four different concrete mixture designs were used to fabricate the ten box beams. Conventional concrete mixtures containing limestone and hard river gravel coarse aggregate were used as well as SCC mixtures containing limestone and hard river gravel coarse aggregate. The use of several different mixture designs along with fifty-five flexural tests provided a complete range of data to recommend an appropriate maximum allowable compressive stress at prestress transfer.

1.3 CHAPTER OUTLINE In Chapter 2 a thorough literature review on four topics concerning the effects of

increasing the allowable compressive stress limit at prestress release is presented. First, the historical background of the allowable stress limit along with recent research pertaining to its increase is presented. Next, the properties and behavior of high strength concrete used in prestressed applications is discussed. The information on high strength concrete provides insight into the behavior of the precompressed tensile zone of the test specimens. Third, the material properties and behavior of self consolidating concrete (SCC) in relation to conventional concrete are presented. The information presented on SCC is pertinent to understanding the behavior of the five box beams cast using SCC. Finally, the two methods used to estimate prestress losses in the test specimens are presented. These two procedures include the NCHRP 496 Detailed Prestress Loss Method (Tadros et al., 2003) and the AASHTO LRFD Interim 2008 Refined Loss of Prestress Estimate (AASHTO, 2008).

In Chapter 3, the fifty-five specimens tested in flexure are described. Forty-five specimens were TxDOT Type-C girders and were fabricated in four different precast plants in the state of Texas. The maximum compressive stress in the Type-C beams ranged from 0.56f’ci to 0.76f’ci. The remaining ten specimens were TxDOT 4B28 box beams and were fabricated by a single fabricator at two different fabrication plants. The maximum compressive stress at prestress transfer for the box beams ranged from 0.64f’ci to 0.66f’ci. Chapter 3 provides the details of the design and fabrication of all fifty five girders as well as the process of storing and shipping the beams to the Ferguson Structural Engineering Laboratory.

In Chapter 4, the experimental setup for all fifty-five test specimens is described. The load protocol, test setup, instrumentation, and data acquisition setup for both Type-C beam flexural tests as well as 4B28 box beam flexural tests is documented. All fifty five test specimens were subjected to statically determinate four point loading with a five foot constant moment region in the center of the beam.

In Chapter 5, the results and analysis of all fifty five flexural beam tests are presented. The experimentally measured cracking load of each Type-C beam and 4B28 box beam specimen was compared to loads predicted using standard design calculations

5

(P/A ± Mc/I). These plots were then used to determine an appropriate allowable compressive stress at release. In addition, observations from five box beams fabricated with SCC are presented.

Finally in Chapter 6, the conclusions from part 1 of phase 2 of TxDOT project 5197 are presented along with a recommendation of increasing the allowable compressive stress at prestress transfer. In addition, recommendations for future research are discussed.

6

7

CHAPTER 2 Literature Review

2.1 OVERVIEW This literature review covers four main topics related to the effects of increasing

the allowable compressive stress limit at prestress release. First, the historical background of establishing the current allowable stress limit, as well as recent research for relaxing this stress limit is presented in Section 2.2. In order to evaluate the adequacy of the current code limit, it is important to understand the origin of the maximum compressive stress limit at prestress transfer as well as the recent research pertaining to increasing that maximum stress limit.

Next, in section 2.3, the properties and behavior of high strength concrete used in prestressed applications are presented. The information presented in this section is pertinent to the behavior of the pre-compressed tensile zone of prestressed girders and therefore, the cracking load of the test specimens. Section 2.3 includes an analysis of the properties and behavior of high strength concrete as compared to those of normal strength concrete. More specifically, the early age strength gain, material properties, response to uniaxial load, and techniques to quantify internal damage are presented.

Third, the properties of self-consolidating concrete (SCC) are compared with those of conventional high strength concrete in Section 2.4. In TxDOT Project 5197, five prestressed 4B28 box girders were fabricated using SCC. An understanding of the material properties and behavior of SCC is needed to effectively evaluate the structural performance of the beams cast with SCC. Section 2.4 contains relevant research pertaining to SCC mixture design, mechanical properties of SCC, and behavior of both small-scale and full-scale girders cast with SCC.

Finally, the two methods used to estimate prestress losses in the test specimens are discussed in Section 2.5. In order to comparatively study the experimentally observed cracking loads among test specimens, time-dependent methods of estimating prestress losses were needed. The methods described in this section provided a consistent and unbiased means for cracking load prediction of the test specimens.

It should be noted that TxDOT Report 5197-1 (2007) contains a thorough literature review regarding similar topics. As such, any information that has already been presented in the previous report (Birrcher and Bayrak, 2007) will be presented here only in summary for the sake of brevity.

2.2 HISTORICAL BACKGROUND OF ALLOWABLE COMPRESSIVE STRESS AT RELEASE In the development of standards and codes for reinforced and prestressed

concrete, the concept of allowable stresses has changed significantly over time. In the following sections, the evolution of allowable stresses for reinforced concrete and prestressed concrete are presented. Additionally, recent research involving the potential

8

relaxation of the current code limit for allowable compressive stress at prestress transfer is presented.

2.2.1 Allowable Stresses in Reinforced Concrete In the early 20th century, a need for code provisions for reinforced concrete was

becoming apparent. In order to provide some guidance for engineers, allowable stresses were established for various stress conditions. The National Association of Cement Users (NACU) in 1910 issued the first document to gain official standing in the United States titled “Standard Building Regulations for the Use of Reinforced Concrete” (Winter 1982). In this publication, the allowable fiber stress for flexural compression was 0.325f’c. Over time, several joint committees along with the American Concrete Institute (ACI) changed these allowable stress limits for reinforced concrete. Table 2-1 presents the development of allowable stresses during the 1900s.

Table 2-1 Development of Allowable Stresses in Reinforced Concrete

Year Document Name Institution Allowable Release Stress 1910 Standard Building

Regulations for Reinforced Concrete

NACU 0.325 f’c for flexural Compression

1916 First Joint Committee Report on Concrete and

Reinforced Concrete

1st Joint Committee on Concrete and

Reinforced Concrete

0.475 f’c for fiber stress in compression adjacent to supports of continuous members. 0.375 f’c everywhere else

1924 Second Joint Committee Report on Concrete and

Reinforced Concrete

2nd Joint Committee on Concrete and

Reinforced Concrete

0.45 f’c for fiber stress in compression adjacent to supports of continuous members. 0.4 f’c everywhere else

1936 ACI 501-36 ACI 0.45 f’c for fiber stress in compression adjacent to supports of continuous members. 0.4 f’c everywhere else

1941 ACI 318-41 ACI 0.45 f’c anywhere along a member After 1941, the allowable stress limit of 0.45f’c remained until ACI accepted

Ultimate Strength Design as the fundamental design approach for reinforced concrete. When this was officially adopted in 1971, there was no need for allowable stresses in traditional reinforced concrete. However, allowable compression limits in prestressed concrete still existed.

2.2.2 Compressive Stresses at Release in Prestressed Concrete In 1942, the American Concrete Institute formed the first committee on

prestressed concrete. This committee investigated design procedures and recommended research for this new use of concrete (Hawkins, 1981). The committee expanded over the next several years to form the joint ACI-ASCE Committee 323 (later 423) on Prestressed Concrete. During this same time period, the Bureau of Public Roads was also developing design recommendations for prestressed concrete. Both entities produced separate

9

documents that specified an allowable stress in compression at prestress transfer of 0.6f’ci for prestressed members and 0.55f’ci for post-tensioned members.

No explicit reasoning for these values was provided. Several authorities on prestressed concrete disagreed over the proper amount of compressive stress that should be allowable at release. Many felt uncertain about specifying such a high stress level, but others felt justified based on empirical practice in the prestressed concrete industry.

In 1961, shortly after these documents were circulated, the American Association of State Highway and Transportation Officials’ (AASHTO) Standard Specifications for Highway Bridges adopted a maximum allowable stress of 0.6 f’ci for prestressed concrete and 0.55 f’ci for post-tensioned concrete. Later in 1963, ACI Committee 318 accepted 0.6 f’ci for all prestressed and post-tensioned construction. These allowable stresses remained unchanged in ACI 318 until very recently and are discussed in section 2.2.3.9.

2.2.3 Recent Research and Discussion There has been significant interest in recent years in the possibility of relaxing the

current allowable compressive stress limit at prestress transfer of 0.6f’ci. There are many potential benefits of relaxing this limit including a faster production schedule for fabricators and producers, the reduction of harped or debonded strands in prestressed concrete beams, and the increase in the number of prestressing strands in a given section, therefore increasing load-carrying capacity. The following sections document the research conducted on this issue and discussion regarding this release limit. As indicated earlier, any information previously presented in TxDOT Report 5197-1 (2007) is presented here only in summary.

2.2.3.1 PCI Standard Design Practice 1996, 1997, and 2003 PCI Standard Design Practice is a document developed by the Prestressed

Concrete Institute (PCI) in an effort to remedy the instances where “ACI provisions are either ambiguous or in conflict with industry practice” (Raths, 2007). The PCI Standard Design Practice was first published in 1996 and amended in 1997 and 2003. All three editions of the document suggest that the current ACI limitation of 0.6f’ci is too conservative. Citing common practice and industry experience, the 2003 edition of PCI Standard Design practice endorses a maximum allowable compressive stress at release of 0.7f’ci.

2.2.3.2 Russell and Pang, 1997 Russell and Pang conducted experimental work on 432 concrete cylinders to

investigate the effect of sustained compressive stress on overall concrete strength. In this study, cylinders were loaded to sustained stress levels of 0.6f’ci, 0.7f’ci, or 0.8f’ci for varying durations. Each cylinder had a companion “control” cylinder that was not loaded. Upon completion of the specified load duration, the cylinders were tested in compression to failure. During this testing, four of the cylinders loaded to 0.8f’ci failed prematurely under the applied compressive stress. Russell and Pang claimed that their

10

research indicated the possibility of relaxing the allowable compressive stress limit to 0.7f’ci. However, they suggested that further research be conducted before making such a change.

The conclusion from Russell and Pang’s work was only based on a few cylinders which failed prematurely. For those cylinders loaded to 0.8f’ci which did not fail prematurely, no significant reduction in strength was noticed. Therefore, analyzing the compressive strength of concrete alone may not be the best means of investigating internal damage in concrete.

2.2.3.3 Huo and Tadros, 1997 Huo and Tadros investigated the effects of early release with the analysis of an

18-inch by 18-inch concentrically prestressed concrete member. Using standard material properties, the researchers ran a linear elastic analysis as well as a separate, iterative non-linear analysis to determine how many strands it would take to crush the concrete upon prestress transfer. Using an f’ci of 3500 psi, the linear elastic method required 45, ½-inch diameter 270 ksi low-relaxation strands. For the non-linear analysis, they assumed that concrete failed when it reached an ultimate strain of 0.003, not when it reached an ultimate stress. Accounting for this fact in their analyses, the non-linear method required 62, ½-inch diameter 270 ksi low-relaxation strands. Huo and Tadros made several observations in this study. First, they showed that both the linear method and the non-linear method provide similar results up to the limitation of 0.6f’ci. Beyond this, the linear method began to underestimate the number of strands required to fail the section. Also, Huo and Tadros made the distinction that a prestressed concrete member differs from a member subjected to external forces because of an internal “self-relieving mechanism” which causes the stress in the strands to decrease over time. Huo and Tadros were not able to make any “definitive recommendations” and recommended further investigation of many other factors such as creep and shrinkage before relaxing the 0.6f’ci compressive stress limit.

2.2.3.4 Noppakunwijai, Tadros, Ma and Mast, 2001 This research advocated the use of a strength design approach in order to evaluate

an appropriate compressive stress at prestress transfer. In this study, the allowable compressive stress at prestress transfer was analyzed as a strength limit state rather than a serviceability limit state. Using a standard PCI rectangular section, the release strength required by ACI 318, PCI, and the proposed strength design method were compared. Additionally, two inverted tee specimens with compressive stresses at release of 0.79f’ci and 0.84f’ci were fabricated and monitored for approximately 100 days. During this time, the researchers were able to predict the change in concrete strain due to shrinkage and creep with reasonable accuracy. Additionally, the initial and long-term camber of the inverted tees was predicted with reasonable accuracy. The authors observed that increased stresses at release cause increased amounts of prestress loss, therefore reducing the stress on the section.

11

At the conclusion of this research, the authors recommended the complete elimination of the allowable compression stress limits in favor of their proposed strength design method.

2.2.3.5 Castro, Kreger, Bayrak, Breen, and Wood, 2004 This project, also referred to as TxDOT project 4086, investigated the effect of

increasing the allowable compressive stress at prestress release. For this study, 30 pretensioned girders were fabricated with compressive stresses at release ranging from 0.46f’ci to 0.91f’ci. The short and long-term camber for these girders was recorded for all specimens and compared to predicted values. Castro et al. (2004) made several observations from this study. First, camber growth increased with higher levels of compressive stress at prestress transfer. Also, the camber at ten days was more accurately predicted for beams meeting the ACI 318 code provisions than beams subjected to higher release stresses. This result indicated the negative impact of subjecting a beam to a higher compressive stress at release. Ultimately, Castro et al. (2004) recommended that increasing the release stress was acceptable “as long as the long-term camber was adequately predicted” (Castro 2004). The results from Castro’s work indicated a need for more research and recommended investigating the live load performance of girders stressed above the current allowable limit. This recommendation initiated TxDOT project 5197, the subject of this report.

2.2.3.6 Hale and Russell, 2006 Hale and Russell investigated the accuracy of prestress losses in concrete bridge

girders stressed above the code allowed limit of 0.6f’ci. For their experimental testing, four I-girders were cast with compressive stresses at release ranging from 0.57f’ci to 0.82f’ci. Prestress losses were measured from concrete surface strains in each of the four girders for one year. All measured prestress losses were accurately predicted within reasonable limits by the available methods used in the research. Additionally, Hale and Russell (2006) observed that the ratio of prestress loss to compressive stress at release was approximately the same for all four bridge girders. Based on this proportionality of prestress loss and compressive stress at release, the authors recommended an increase in the allowable release stress from 0.6f’ci to 0.7f’ci.

It is important to note that although they only recommended increasing the code limitation to 0.7f’ci, their reasoning would justify an increase in the allowable stress limit to 0.82f’ci. Additionally, nonlinear behavior of the prestressed girders was taken into account to accurately estimate deformations at prestress transfer. This practice is not common in prestressed concrete design.

2.2.3.7 Dolan and Krohn, 2007 The investigation by Dolan and Krohn consisted of a thorough literature review

on compressive stresses at prestress transfer and a survey of the professional and producer members of PCI. This work was done in an effort to “determine the current

12

states of research and practice regarding compression transfer stress of prestressed concrete members” (Dolan and Krohn 2007). The electronic survey was completed by 61 respondents. Of those respondents, 36 stated that they “routinely use compression transfer stresses above 0.6f’ci” (Dolan and Krohn 2007). Figure 2-1 displays the magnitude of compression transfer stresses used by respondents of the electronic survey.

Figure 2-1 Compressive Stress at Transfer used by PCI survey respondents (adopted

from Dolan and Krohn, 2007) Also during the survey, respondents were given the opportunity to identify any

problems associated with the use of an increased allowable stress at prestress transfer. Some problems such as excessive camber or concrete splitting at the ends of the members were noticed and reported. However, a large majority of the respondents noted no problems with using high stresses at release. After evaluating the surveys and investigating literature, Dolan and Krohn concluded that producers have many years of experience in fabricating prestressed concrete beams using compression stresses at release greater than 0.6f’ci with minimal problems. Therefore, they recommended that ACI 318 should raise the allowable compressive stress at prestress transfer to 0.7f’ci. In addition, they recommended that the PCI equations for prestress loss and camber be reviewed and revised for this increased transfer stress.

0

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0.5 0.55 0.6 0.65 0.68 0.7 0.75 0.8

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13

2.2.3.8 Birrcher and Bayrak, 2007 Following the research of Castro et al. (2004), TxDOT project 5197 was initiated

to evaluate the live load performance of prestressed concrete girders stressed beyond the allowable limit. This research study involved the static live load testing of 24 scaled girders from Castro et al. and 12 full-scale TxDOT standard A-beams and the fatigue testing of 4 beams from Castro et al.

The static live load tests were used to experimentally determine the cracking load of the test specimens. These experimentally determined cracking loads were then compared to analytically predicted cracking loads. The following three methods were used to analytically determine cracking loads: PCI Design Handbook Loss of Prestress Estimate (PCI, 2004), the NCHRP Report 496 Detailed Prestress Loss Method (Tadros et al., 2003), and the AASHTO LRFD Refined Loss of Prestress Estimate (AASHTO, Interim 2005).

During the static tests, the specimens were subjected to four-point loading creating a constant moment region at the center of the beam. While testing the specimens, loading was stopped several times to mark cracks as well as document crack widths and crack propagation. All test specimens were loaded until significant cracking was observed. Upon completion of these static tests, the accuracy of the predicted cracking load was plotted against the maximum compressive stress at release. These plots are displayed in Chapter 1 and Chapter 5 of this report. From these plots, Birrcher and Bayrak observed an unconservative drop in cracking load accuracy with beams subjected to a maximum compressive stress higher than 0.65f’ci.

The results from this work indicate that “an increase of the allowable compressive stress to 0.65f’ci was justified” (Birrcher and Bayrak, 2007). Although Birrcher and Bayrak did believe an increase to 0.65f’ci was justified, they recommended further testing of beams with different shapes, different fabricators, and different concrete mixture designs. The recommendations from Birrcher and Bayrak are the basis for the research presented in this report. The complete documentation of their research is located in a Center for Transportation Research (CTR) report under TxDOT Project 5197-1.

2.2.3.9 Summary of Recent Research Table 2-2 provides a summary of the recent research regarding compressive

stresses at prestress transfer. Much of the research in recent years has supported changing the allowable release factor to 0.7f’ci with only a few concerns about girder performance. Therefore, in 2008, ACI 318 adopted a new allowable compressive stress limit of 0.7 f’ci at the ends of prestressed members citing, “research in the precast, prestressed concrete industry practice” (ACI 2008). ACI maintained the limit of 0.6 f’ci at midspan. However, the Interim 2008 Edition of the American Association of State Highway and Transportation Officials (AASHTO-LRFD) Bridge Design Specification maintained the allowable compressive stress at transfer of 0.6f’ci everywhere. These limitations are current as of the date of this report.

14

Table 2-2 Summary of Recent Work Regarding Compressive Stress at Release

Researchers Year Focus of Research (In regard to Compressive Stress at Release)

Scope of Experimental Work

Russell and Pang 1997 Compressive Strength 432 cylinders Huo and Tadros 1997 Nonlinear Behavior None Noppakunwijai et al. 2001 Creep, Shrinkage, Camber, etc. 2 IT girders Castro et al. 2004 Camber 30 Rect., IT, T girders Hale and Russell 2006 Effective Prestressing Force 4 I-Beams Dolan and Krohn 2007 Current State of the Industry None

Birrcher and Bayrak 2007 Live Load Performance 12 Full Scale I-Beams 24 Scaled girders

2.3 MECHANICAL PROPERTIES AND BEHAVIOR OF HIGH STRENGTH CONCRETE In order to produce beams efficiently, it has been common practice to use high-

strength concrete (HSC) to fabricate prestressed concrete girders. Research on this topic has shown that high strength concrete has substantially different mechanical properties than normal strength concrete. Figure 2-2 displays the theoretical stress-strain responses for several concrete strengths. This figure was generated with expressions for stress and strain developed by Thorenfeldt, Tomaszewicz, and Jensen (1987) found in Chapter 3 of the Collins and Mitchell text, Prestressed Concrete Structures (1997). As shown in Figure 2-2, the ascending and descending branches of the stress-strain curves become much more linear as the concrete strength increases. Figure 2-2 also illustrates that the response of low strength concrete is substantially different from that of high strength concrete. In addition to stress-strain behavior, the magnitude of shrinkage and creep differs between normal strength and high strength concrete. It is for these reasons that equations specific to high strength concrete have been developed to calculate the modulus of elasticity, shrinkage strains, creep strains, and other mechanical properties.

15

Figure 2-2 Theoretical stress vs. strain response for various concrete strengths

This section details the high strength concrete research on early age mechanical properties, response to uniaxial loading, and methods to measure internal damage. A clear understanding of the material properties and behavior of high strength concrete is needed to evaluate the behavior of the pre-compressed tensile zone of prestressed concrete bridge girders.

2.3.1 Mechanical Properties of High Strength Concrete at Early Ages In the prestressed concrete beam fabrication industry, high strength concrete is

often loaded to large stresses at early ages (less than 24 hours after casting). Therefore, it is important to study and understand the early age strength and behavior of high strength concrete. The 28-day properties of high strength concretes have been extensively researched, but few investigations have focused on the properties of high strength concrete at younger ages. The following section highlights work on the early age mechanical properties of high strength concrete.

2.3.1.1 Khan, Cook, and Mitchell, 1995 This research study includes an investigation of the stress-strain behavior of three

different types of concrete: Low Strength Concrete, LSC, at 4000 psi (30 MPa), Normal Strength Concrete, NSC, at10,000 psi (70 MPa) and High Strength Concrete, HSC, at

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16

14,500 psi (100 MPa). Concrete cylinders were batched and tested in compression during the first 72 hours after batching in order to investigate the concrete strength with temperature rise, stress-strain behavior at early ages, and influence of curing conditions on compressive strength and elastic modulus. The researchers concluded that for all three types of concrete (LSC, NSC, HSC) there is, “a significant difference in the shape of the compressive stress-strain response at a very early age” (Khan et al., 1995). Figure 2-3 displays the stress strain behavior of a 10,000 psi (70 MPa) concrete at various ages. It is important to note that the behavior at 16.5 hours is much more nonlinear than at 3 days. As the compressive strength increases, the ascending and descending branches of the stress-strain curve become much more linear.

Figure 2-3 Stress-Strain behavior for a 10,000-psi concrete at early ages (Khan et al.,

1995 adopted from Birrcher and Bayrak, 2007) The nonlinearity of concrete at a young age as shown in Figure 2-3 indicates that

significant internal damage may occur if the concrete is loaded to high levels of compression during the first 24 hours after casting.

2.3.2 Response to Uniaxial Loading and Quantifying Internal Damage Concrete is made up of cement paste and aggregates. Both of these materials

individually have linear and highly brittle stress strain curves (MacGregor and Wight, 2005). When combined, the stress-strain behavior becomes nonlinear. This nonlinear behavior can be described by the phenomenon of microcracking. Microcracks are tiny internal cracks 1/8-inch to ½-inch in length. Under increasing compressive stresses, the microcracks expand causing nonlinearity and eventually cause failure of concrete in

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17

compression. The following research studies document the formation and propagation of this internal damage.

2.3.2.1 Richart, Brandtzaeg, and Brown, 1929 An investigation of the testing of short columns (40-inches tall and 10-inches in

diameter) that were either plain concrete or spirally reinforced was presented in this research. These short columns were tested in compression in an effort to understand the “action of concrete under compressive stresses” (Richart et al., 1929). Upon testing the plain concrete columns with no reinforcing bars, the researchers deemed it appropriate to divide the concrete response into three stages. In the first stage, the behavior was elastic with stresses and strains proportional to each other. This first stage was observed for at least 25-percent of the ultimate load. The second stage was defined with appreciable deviations from the previous stress-strain proportionality. Finally, the third stage was defined by an abrupt change in lateral deformation of the specimen. During the third stage, the internal structure broke down leading to ultimate failure. This third stage occurred at 75 to 85-percent of the ultimate load. A visual depiction of these three stages is displayed in Figure 2-4.

Figure 2-4 Depiction of three stages discussed by Richart et al. 1929, adopted from

Birrcher, 2007 Richart et al. concluded that the phenomenon observed in the third stage was due

to “internal splitting or breaking of the continuity of the material” (Richart et al., 1929). As such, serious damage to the microstructure was observed when concrete was loaded to high stresses (about 75-percent of ultimate strength).

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18

2.3.2.2 Hsu, Slate, Sturman, and Winter, 1963 The type and extent of microcracking in plain concrete cylinders subjected to

uniaxial compressive loads was investigated in this research project. Ten representative specimens were examined and reported. The cylinders were loaded to a prescribed amount of strain between 0 and 0.003. After loading, the cylinders were cut with a diamond blade masonry saw and examined under a microscope. Using ink dye to accentuate the cracks, the researchers extensively analyzed the specimens and drew cracking maps. Hsu et al. determined that three types of microcracks exist: bond cracks, mortar cracks, and aggregate cracks. Bond cracks exist at the interface between the aggregate and mortar and are present before loading even begins. After loading is initiated, these bond cracks increase in size and number. As bond cracks propagate, stress in the concrete must redistribute causing cracks within the mortar to develop. Upon reaching 70 to 90-percent of the ultimate load, the mortar cracks increase and bridge between bond cracks to form continuous crack patterns. This formation of continuous crack patterns signifies a “critical” load at which the stress-strain diagram becomes highly non-linear indicating the beginning of the breakdown of the internal structure.

Essentially, Hsu et al. (1963) agreed with the findings of Richart Brandtzaeg and Brown (1929). Both Hsu et al. and Richart et al. indicated critical loads of 0.7Pult and 0.75Pult respectively where the internal structure of the concrete begins to break down. It is important to note that relatively low strength concrete samples were used in both studies. Even though the compressive strengths investigated by Hsu et al. and Richart et al. were low, the same inelastic behavior exists for normal-strength concrete today (MacGregor, 2005).

2.3.2.3 Ngab, Slate, and Nilson, 1981 Ngab et al. analyzed the microcracking in eighty-four 3.5 x 3.5 x 10-inch

specimens of normal strength and high strength concrete subjected to compressive loading. Specimens were tested for different durations at different ages after curing. The applied stress-to-strength ratio of the specimens varied from 0.30 to 0.85. After loading, the specimens were cut with a diamond-blade masonry saw and analyzed. After digitizing, measuring, and mapping the microcracking present in the concrete, the researchers made several observations: First, the amount of microcracking in high strength concrete was significantly less than that of normal strength concrete. Additionally, the amount of microcracking for both high strength and normal strength concrete was found to be a function of the strain imposed on the concrete, regardless of whether the strain was due to short-term loads, sustained loads, or shrinkage.

In addition to analyzing microcracking in concrete, Ngab et al. investigated the creep behavior of these specimens. The researchers concluded that for high strength concrete, the ratio of applied stress and creep deformation was linear until approximately 70-percent of f’c. Based on these findings, it was concluded that additional creep

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deformation, and therefore microcracking, could develop at or near loads of 70-percent of f’c.

2.3.2.4 Smadi, Slate, and Nilson, 1985 and 1987 These two research investigations focused on the sustained loading response of

high-strength concrete relative to medium- and low- strength concrete. In this study, high-strength concrete was defined as concrete with compressive strengths between 8500 psi and 10000 psi. All test specimens consisted of 4 x 8-inch cylinders loaded to a specific level of stress ranging from 0.4f’c to 0.95f’c. Many specimens loaded to very high values of stress failed under the applied load. Specimens which did not fail were monitored for creep deformation. From this research, Smadi found that high strength concrete achieves 80- to 85-percent of its short term compressive strength when placed under long-term sustained loading. For the specimens that did not fail, Smadi also found that high strength concrete performed better than medium- and low-strength concrete in regards to creep deformation. Furthermore, creep strain appeared to increase linearly with applied stress until the load reached approximately 70-percent of the ultimate capacity. From that point forward, the creep deformation became nonlinear. Smadi concluded that the “deviation of creep from linearity with increasing stress is believed to be due to a significant increase in bond cracks along the mortar-aggregate interface” (Smadi et al., 1987). Smadi et al. also recommended that “the stresses in HSC can be increased safely up to the creep proportional limit, or up to about 65 percent of ultimate [capacity], without causing significant crack formation” (Smadi et al., 1987). These results suggest that loading in a range beyond 0.65f’ci could cause significant cracking and elevated creep deformation in high strength concrete.

2.3.2.5 Delibes Liniers, 1987 Delibes Liniers investigated the loss of tensile strength in concrete after axial

compressive loading. Liniers believed that the internal microcracking in the concrete weakened the tensile strength. The experimental testing involved casting concrete cylinders and loading them in compression to a specified value of their ultimate strength (50-percent to 95-percent) and maintaining the load for a specified period of time. After loading, the cylinders were split in accordance with ASTM C496-71. Delibes Liniers found that tensile strength decreased with increasing compressive stress to strength ratios. Figure 2-5 shows the summarized results from this research investigation.

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Figure 2-5 Tensile strength loss as a function of compressive stress for various

compression times and general curing conditions (Delibes Liniers, 1987). As shown in Figure 2-5, loading at higher compressive stress-to-strength ratios

lead to increased losses in tensile strength. It is important to note that Delibes Liniers found that microcracks typically form parallel to the applied load. Liniers notes that “ ‘damage’ in a direction perpendicular to the compression was almost negligible” (Delibes Liniers 1987). This finding is interesting because research by Ngab et al. and Smadi et al. indicated microcracks perpendicular to the applied load.

Ultimately, Delibes Liniers concluded that “the necessity of limiting compressive stress under 60% of the strength is confirmed” (Delibes Liniers, 1987). This research showed that tensile strength loss was a reasonable indicator for internal damage and that concrete loaded above 60-percent of its strength could experience tensile strength loss and therefore significant internal damage.

2.3.2.6 Gettu, Aguando, and Oliveira, 1996 Gettu et al. investigated internal damage in concrete induced by monotonic and

cyclic compressive loading. In order to evaluate the amount of damage in concrete specimens, splitting tension tests were performed. For the experimental research, cubes were cast, cured, and subjected to either monotonic or cyclic loads. After loading, the cube was turned 90-degrees about the casting direction and a splitting tension test was performed. This summary only contains the conclusions that were based on the data for the specimens loaded monotonically.

Once the cubes had cured for 28 days, they were loaded to a specified applied stress-to-strength ratio ranging from 0.25 to 0.85. After maintaining the load for fifteen minutes, the specimen was split in tension so that the failure plane was parallel to the

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previously applied compressive stress. Loading in this fashion was analogous to the testing done by Delibes Liniers (1987). Upon completing this experimental research, Gettu et al. concluded that “Damage in high-strength concrete due to uniaxial compressive stress is negligible as long as the monotonically applied stress is less than 60 percent of its strength” (Gettu et al., 1996). In summary, the findings of Gettu et al. (1996) coincide with Delibes Liniers (1987). Both investigations showed that loss in tensile strength provided a good indicator of internal damage. In addition, both research experiments indicated that loading concrete specimens to compressive stresses above 60-percent of their ultimate strength cause internal damage and loss of tensile strength.

2.4 PROPERTIES OF SELF-CONSOLIDATING CONCRETE In the current research project (TxDOT Project 5197), five box beams were

fabricated using self consolidating concrete (SCC). In order to adequately evaluate the behavior of these beams, a thorough understanding of SCC is needed. The following section includes relevant research pertaining to SCC mixture design, mechanical properties of SCC, and behavior of full-scale girders cast with SCC.

Self consolidating concrete (SCC) is a type of concrete that flows freely to fill formwork and requires little to no external vibration. In order to obtain these unique characteristics for use in prestressed concrete, several changes must be made to conventional high strength concrete mixture designs. First, the concrete must become more fluid to adequately fill formwork and congested areas of reinforcing steel without mechanical vibration. High fluidity in concrete is achieved by using a plasticizing admixture, increasing the water to cementitious materials (w/cm) ratio, using a smaller coarse aggregate size, or a combination of the previous three. By making the concrete more fluid, the mixture is much more likely to segregate during casting, creating a weak and vulnerable layer of cement paste at the top of the vertically cast members. In order to combat segregation, the ratio of fine aggregate to coarse aggregate can be increased, the w/cm ratio can be decreased, and/or a viscosity modifying admixture (VMA) can be used. All of these changes help to suspend the coarse aggregate in the mixture and prevent segregation.

Although the composition of SCC mixtures can vary greatly, two main design methodologies have developed (Bonen, 2005). The first design method is to utilize superplasticizer, low w/cm, and a high amount of fine aggregate. This method is often referred to as the “powder” method because of the relatively large amount of the cementitious materials and fine aggregate in the mixture. The second design method is to incorporate a viscosity modifying admixture. Adding the VMA to the mixture will control the concrete viscosity and therefore reduce the amount of powder needed in the mixture. A visual depiction of the two types of self consolidating concrete as compared to ordinary concrete is displayed in Figure 2-6.

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Figure 2-6 Typical volume fraction for different types of concrete (adopted from

Bonen, 2005). The use of self consolidating concrete has many potential benefits for prestressed

applications. First, the ability of SCC to flow freely to fill formwork without vibration provides a savings in time and labor. With conventional concrete, the mixture must be properly vibrated by laborers to fill the formwork with layers of concrete free of voids or defects. This can become time-consuming and troublesome in areas of highly congested reinforcing steel or complex geometry. SCC mixtures provide a simpler means of placing concrete. Additionally, due to the fluid nature of the mixture, SCC is less likely to have void-related imperfections such as bug holes, honeycombs, etc. In theory, this provides higher quality concrete requiring less repair.

Although SCC offers many potential benefits to precast beam fabricators, there are some concerns which must be investigated. Because the SCC is very fluid, there is a constant concern regarding segregation. If the mixture does not adequately suspend the coarse aggregate, then the coarse aggregate will sink, leaving behind weak layers of cement paste and a non-uniform structural material. In addition, D’Ambrosia et al. (2005) has shown that SCC mixtures exhibit high amounts of shrinkage and a lower modulus of elasticity when compared to conventional concrete. Furthermore, the “raw materials cost of SCC is higher by about 13-30% than that of conventional mixtures with similar mechanical properties” (Bonen, 2005). This additional cost can be mitigated by the savings from labor and time, but that may not always be the case.

0%

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Ordinary Concrete SCC VMA Type (FA+SF+VMA)

SCC Powder Type

Volu

me,

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Volume Fraction with 6% nominal air

Coarse Aggregate

Sand

Cementitious

Water

Air

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Many departments of transportation around the country have been reluctant to adopt the use of SCC until adequate research assured that the use of SCC would provide acceptable results. The following sections detail recent research relevant to the mechanical properties of SCC and applications of SCC in prestressed concrete.

2.4.1 D’Ambrosia, Lange, and Brinks, 2005 D’Ambrosia et al. investigated the performance of self consolidating concrete at

early ages (less than 24 hours). The research team monitored the early age shrinkage, creep, strength gain and modulus of elasticity in an effort to assess the characteristic behavior of SCC. In addition, an effort was made to understand how different mixture design strategies would impact early age properties of hardened concrete.

First, a database of more than 150 SCC mixtures was compiled in an effort to understand SCC mixture proportioning. Upon analysis of the data, the researchers found that SCC mixtures trended toward lower water-to-cementitious material ratios (w/cm) when compared to normal strength concrete. Also, the aggregate content tended to be lower for SCC when compared to typical concrete. Furthermore, SCC mixtures typically had a higher proportion of sand to provide segregation resistance.

After analyzing typical SCC mixtures, four different mixture designs were selected for testing. A conventional high strength mixture design was used as a control and given the designation OPC 1. The proportions for all five mixtures are given in Table 2-3.

Table 2-3 Concrete mixture proportions (D’Ambrosia, 2005) Material OPC 1 SCC 1 SCC 2 SCC 3 SCC 4 Portland Cement (lb/yd3) 726 661 601 685 679 Fly Ash (lb/yd3) 0 157 325 0 151 Coarse Aggregate: 3/4” maximum (lb/yd3) 1853 367 1365 1627 579 Coarse Aggregate: 3/8” maximum (lb/yd3) 0 1075 0 0 1018 Fine Aggregate (lb/yd3) 1192 1403 1336 1389 1389 Water (lb/yd3) 290 311 301 278 267 Superplasticizer (fl oz/yd3) 22 63 29 49 36 VMA (fl oz/yd3) 0 0 0 22 0 w/cm 0.4 0.38 0.33 0.41 0.34

Specimens from all the mixture designs in Table 2-3 were tested for early age

strength, and autogenous and drying shrinkage. To monitor the early age strength gain, two sets of 4 x 8-inch cylinders were subjected to compressive strength tests according to ASTM C39. One set of cylinders was kept in a temperature and humidity controlled environment while the other set was moist cured. In addition, early age indirect tensile strength was measured on the temperature and humidity controlled cylinders in accordance with ASTM C496. Autogenous shrinkage was evaluated with embedded strain gauges on 3 x 3 x 11-inch prisms in a sealed container. The temperature of the specimens was monitored so thermal expansion could be taken into account. To measure restrained and unrestrained drying shrinkage, a uniaxial test method was employed. This

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involved two companion 3 x 3 x 24.5-inch specimens that were cast and allowed to cure under a plastic sheet for one day. The unrestrained shrinkage specimens were used to monitor free shrinkage measured by using a linear variable differential transducer (LVDT). The restrained shrinkage specimens were placed in a testing machine that applied the necessary tensile force to maintain the specimen’s original length.

The compressive strength development testing indicated high early strengths due to the tendency for SCC mixtures to have a low w/cm ratio and high paste content. In order to maintain a similar curing regime, the concrete cylinders tested for tensile strength were kept in a drying environment with the creep and shrinkage specimens. The results from the tension tests indicated that the tensile strength did not increase very much after three days because the concrete was not given an external source of moisture. This led to the conclusion that the “lack of external water during curing increased the risk for cracking at early age by limiting strength development” (D’Ambrosia et al., 2005).

The shrinkage portion of the study indicated that SCC had high amounts of drying shrinkage. The results from the total shrinkage of the concrete prisms are displayed in Figure 2-7.

Figure 2-7 Total Shrinkage of 3 x 3 x 24.5-inch Specimens (D’Ambrosia et al., 2005)

As is displayed in Figure 2-7, the drying shrinkage for all four SCC mixtures was noticeably larger than the ordinary concrete mixture. Significant drying shrinkage occurred with SCC when the w/cm ratio was 0.4 or lower. The researchers noted that, “Low w/cm and high paste content caused a significant amount of autogenous shrinkage,

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-10000 5 10 15 20 25 30

AGE (Days)

FREE

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10-6

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which was a major contributor to overall early age shrinkage and subsequent stress development and cracking” (D’Ambrosia et al., 2005).

At the conclusion of their research, the authors recognized a “potentially high risk for cracking” and noted that the low w/cm ratios and high cement paste contents found with SCC mixtures generally contributed to a higher early age cracking risk. The researchers also pointed out that SCC typically had a high early strength gain which could mitigate some of this cracking risk. Ultimately, the researchers reported that different methods of mixture proportioning led to SCC materials with similar flow characteristics, but different hardened mechanical properties.

2.4.2 Ozyildirim, and Lane, 2003, and Ozyildirim, 2005 Ozyildirim and Lane documented an investigation on self consolidating concrete

conducted by the Virginia Department of Transportation (VDOT). Their research report summarizes the early development and testing of SCC mixtures for use by VDOT. In addition, the report describes the use of SCC in VDOT applications.

In developing suitable SCC mixtures, VDOT prepared fifteen concrete mixtures and tested them for consistency, compressive strength, permeability, drying shrinkage, air voids, and freeze-thaw resistance. All tests performed on the hardened concrete samples and the corresponding specifications are presented in Table 2-4.

Table 2-4 Hardened Concrete Tests and Specifications

Tests Specification Age (days) Size (in) Compressive Strength AASHTO T 22 a 4 x 8 Permeability AASHTO T 277 28b 2 x 4 Drying Shrinkage ASTM C 157 28 3 x 3 x 11¼ Freeze-Thaw Analysis ASTM C 666 c 3 x 4 x 16 Air Void Analysis ASTM C 457 28 4 x 8 aAt 28 days for lab specimens and 1, 7, and 28 days for plant specimens. bCured 1 week at 73°F (23°C) and 3 weeks at 100°F (38°C). cSpecimens cured 2 weeks moist and air dried 1 week before testing. Test water contained 2% NaCl.

The slump flow and U-tube (or U-box) tests were used to evaluate workability and consistency. Figure 2-8 displays a picture of the slump flow test and Figure 2-9 displays a schematic diagram of the U-tube test. In the slump flow test, concrete is poured into an inverted slump cone. The slump cone is then lifted allowing the concrete to flow out. The diameter of the circular spread of the concrete provides a measure of flow and therefore determines the consistency of the concrete. In the U-tube test, a “U”-shaped testing apparatus is used (Figure 2-9). This apparatus consists of two legs separated by a wall. The wall extends for most of the “U” except at the bottom where the wall is replaced with reinforcing bars. First, a gate is lowered over the reinforcing bars preventing any flow between legs of the “U”. Then, concrete is poured to the full height of one leg of the “U”. Finally, the gate is removed so the SCC flows through the rebar and rises up the other leg of the “U”. The distance that the concrete rises provides a measure of workability.

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Figure 2-8 Slump Flow Test (TxDOT Research Study 5197)

Figure 2-9 Schematic Diagram of a U-Tube (U-Box) Test (Khayat et al., 2004)

During the slump flow and U-tube tests, it was determined that many of the batches had segregation problems and that, “segregation was an issue that must be watched closely” (Ozyildirim, 2005). Although many SCC mixture designs did not have adequate flow properties, the researchers reported that some mixtures were “viable candidates for SCC” (Ozyildirim, 2005).

In addition to laboratory testing, VDOT had two precast plants batch two different SCC mixtures. These samples were tested in the same manner as the laboratory batches. With the precast plant mixtures, no segregation or bleeding was found and the air contents of the mixtures were within allowable limits. However, shrinkage values of the

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concrete were greater than the recommended maximum value of 400 microstrain at 28 days. Additionally, a poor air-void system was found within the hardened concrete. These tests results indicated that high shrinkage and poor freeze-thaw resistance can occur with SCC mixtures (Ozyilidirim, 2005).

2.4.3 Ozyildirim and Davis, 2007 This research presented an evaluation of the use of SCC in full-scale bulb-tee

beams used in a bridge on Route 33 in Virginia. Ozyildirim and Davis evaluated the fabrication and placement of SCC concrete and monitored strain and camber over time using vibrating wire gauges (VWG).

Prior to placing SCC beams in an actual bridge, two full-scale specimens were cast and tested for transfer and development length as well as shear and flexural capacity. The test beams’ cross section was similar to the cross section of the actual bulb-tee girders in the Route 33 bridge. The beams were 45-inches tall and 60-feet long. Prior to testing, these beams were topped with a conventional concrete slab. In the laboratory, the beams were loaded in four combinations of moment and shear in order to fail the beams in various different modes including flexure, shear, and flexure-shear. No specific information was given in regards to the location of loading, but the researchers reported that, “a strand development failure was never realized” and “the strands were well bonded to the concrete near the beam end, and strand slip was minimal” (Ozyildirim and Davis, 2007). Furthermore, the experimental testing indicated that “the tensile capacity of the concrete in the beam web was 5.3 times the square root of the compressive strength” (Ozyilidirim and Davis, 2007). The value reported for tensile capacity was significantly lower than values typically used in flexural design (6 . or 7.5 ). In regards to the transfer lengths, the measured values were less than predicted by theoretical calculations. Although this observation was reported, there was no explanation of what theoretical calculations were used and the test method to evaluate the transfer length. The researchers concluded that “the test beams behaved at least as well as would be expected for normally consolidated concrete beams” (Ozyildirim and Davis, 2007).

Based on the previous testing of the bulb-tee beams, full scale SCC bridge beams were fabricated and placed in the Pamunkey Bridge on Route 33 in Virginia. For research purposes, eight SCC beams were used to build one of the 49 spans of the bridge. A “control” span of conventional concrete beams was located immediately next the SCC span. Two of SCC beams and two of the conventional concrete beams contained vibrating wire gauges (VWG) to monitor the behavior during service conditions and thermocouples to monitor the beam temperature while curing. The VWGs were installed near the strands in the top and bottom flanges of the beams during fabrication and the thermocouples were placed at mid-depth in the web. All beams and corresponding test samples were steam cured. Table 2-5 displays the hardened concrete properties of the four instrumented bridge beams evaluated in this study.

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Table 2-5 Hardened Properties of Instrumented Bridge Beams (Ozyildirim and Davis, 2007)

Property Age B1 (SCC) B2 (SCC) B3 (control) B4 (control)

Compressive Strength (psi)

2 days 7,470 6,650 6,270 5,790 7 days 9,1790 8,860 7,760 6,960 28 days 10,110 10,700 7,960 7,610 1 year 11,230 10,940 9,750 8,730

Elastic Modulus (ksi)

2 days 5.07 4.54 4.99 4.52 7 days 5.10 5.06 5.45 5.15 14 days 5.00 5.19 5.69 5.16 28 days 4.86 5.35 5.26 4.98 1 year 5.44 5.16 5.80 5.33

Splitting Tensile Strength (psi)

7 days 760 695 715 650 28 days 820 755 675 565 1 year 840 895 805 810

Permeability (coulombs) 28 days 869 996 1,011 985 Shrinkage (microstrain) 112 days 295 255 328 320

As is shown in Table 2-5, the SCC batches obtained higher concrete strengths and

lower shrinkage values than did the conventional concrete mixtures. However, the elastic modulus values for SCC were equal or lower than that of regular concrete. In addition, camber measurements were recorded for the instrumented beams. The camber in the SCC beams was “slightly higher than that in the conventional concrete at an early age, but cambers and strains for all beams have been similar in service” (Ozyildirim and Davis, 2007). The researchers reported that further monitoring of the long-term shrinkage and creep of these beams was underway and the results would be reported when a sufficient number of seasonal cycles have passed.

After fabrication, monitoring, and testing of these bulb-tee beams, the researchers concluded that “SCC members can be designed by using the same methods, assumptions, and limiting values as used for normally consolidated concrete” (Ozyildirim, 2007). However, it was also reported that the mixture proportions of SCC must be carefully monitored because SCC was sensitive to water content. Additionally the researchers reported that, low slump flows (small diameter readings) from the slump flow test leads to concrete which may not self-consolidate. Conversely, high slump flows (large diameter readings) from the slump flow test may lead to segregation problems.

2.4.4 Gross, Yost, and Gaynor, 2007 Gross et al. investigated the time-dependent behavior of prestressed beams cast

with self consolidating concrete. In the project, four T-beams were cast using conventional high-strength concrete (HSC) and four separate T-beams were cast with SCC. Each beam was 124-inches long and 5.5-inches deep. Each beam was prestressed with a single 0.208-inch diameter deformed high-strength steel wire located 0.854-inches above the bottom surface of the beam.

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To monitor time-dependent behavior, strain gauges were attached to each specimen. One resistance bonded strain gauge (standard foil gauge) was attached to the prestressing strand at midspan prior to stressing. The remaining instrumentation consisted of concrete surface strain gauges applied at midspan at the locations specified in Figure 2-10. The dimensions and strain gauging locations for each of the beams is displayed in Figure 2-10.

Figure 2-10 Beam Cross section and strain gauge locations (adopted from Gross et al.,

2007) The mixture proportions used in the conventional high-strength concrete (HSC)

and self-consolidating concrete (SCC) are included in Table 2-6. Both the SCC and HSC mixtures were designed to reach a 2 day compressive strength of 6000-psi and a 28 day compressive strength of 8000-psi.

Table 2-6 Concrete Mixture Proportions per cubic yard (Gross et al., 2007)

Material HSC SCC Coarse Aggregate 1798 lb 1207 lb

Fine Aggregate 1175 lb 1254 lb Water 300 lb 370 lb

Type III Cement 726 lb 983 lb Silica Fume 63 lb 74 lb

High Range Water Reducer 99 oz 180-191 oz Viscosity Modifying Admixture 0 oz 21 oz

w/cm 0.38 0.35

The eight T-beams in this study were cast in pairs. From each pair, one beam was subjected to a sustained load of 530-pounds at 29 days (Beam B) while the other beam remained unloaded (Beam A). Thus, four beams (B Beams) were subjected to sustained loads while the remaining four were not loaded (A Beams). Companion 4-x8-inch cylinders were cast with each of the beams in this study. The companion cylinders were

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used to determine the compressive strength and elastic modulus of the concrete in each specimen. The average 28-day compressive strength of the SCC and HSC mixtures was 8640-psi and 8200-psi respectively. The average measured 28-day modulus of elasticity for the SCC and HSC mixtures was 5140-ksi and 6600-ksi respectively. It is important to note that the 28-day elastic modulus of SCC mixture was significantly lower than that of the HSC mixture.

Using the applied strain gauges, camber and prestress loss were monitored for 300-days. Figure 2-11 displays a plot of measured prestress loss with time for beams cast with SCC and HSC.

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Figure 2-11 Measured prestress loss versus time for SCC beams and HSC beams,

respectively (Gross et al., 2007) As is shown in Figure 2-11, the SCC beams exhibited significantly more prestress

losses than did the beams cast with HSC. As was expected, the beams with sustained service loads (B Beams) exhibited less prestress loss due to the elastic gain imposed by the placement of the dead load. In addition to prestress loss, camber was also measured over the 300-day period. Results of the camber data showed that the beams with SCC exhibited higher cambers than their counterparts cast with HSC.

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From these results Gross et al. concluded that, “Measured prestress losses and net camber after 300 days were significantly higher in the SCC beams as compared to the HSC beams” (Gross et al., 2007). They conjectured that the higher losses and cambers in SCC beams “can have a marked impact on the serviceability behavior of self-consolidating concrete prestressed beams” (Gross et al., 2007). It is important to note that the tests performed by Gross et al. were on a small-scale specimens and may not necessarily represent the response of a full-scale specimen.

2.4.5 Ruiz, Staton, Do, and Hale, 2007 Similar to the experimental research program of Gross et al. (2007), this study

provided a comparison of prestress losses in beams cast with SCC to those cast with conventional concrete. In this research program, 20 prestressed beams were cast and prestress losses were measured on 10 beams. Of those 10 beams, 3 were cast with a conventional concrete mixture (HSC), 2 were cast with an SCC mixture containing Type III cement (SCCIII), and 4 were cast with an SCC mixture containing Type I cement (SCCI). All three mixtures were designed to have an initial release strength of 7000-psi and a 28 day strength of 12,000-psi. The actual average release strength (at 24-30 hours) for the HSC and SCC mixtures was 9260-psi and 7760-psi respectively. The actual average 28-day strength for the HSC and SCC mixtures was 12,100-psi and 11,530-psi respectively. The mixture designs for all three types of concretes are included in Table 2-7.

Table 2-7 Concrete Mixture Designs (Ruiz et al., 2007)

Material SCCI SCCIII HSC Cement (lb/yd3) 950 808 900 Fly Ash (lb/yd3) 0 142 0

Coarse Aggregate (lb/yd3) 1350 1350 1800 Fine Aggregate (lb/yd3) 1474 1400 1207

Water (lb/yd3) 285 304 234 w/cm 0.30 0.32 0.26

HRWR1 (fl.oz/cwt) 7.8-14.5a 8.0-9.0a 9.0-10.5a HRWR2 (fl. oz/cwt) 0-3.0a 0 0 VMA (fl. oz/cu. yd) 0-30.4a 0-30.4a 0

a Dosages of admixtures varied due to variations in ambient air temperatures during time of batching for individual mixtures

HRWR= High Range Water Reducer VMA= Viscosity Modifying Admixture

The beams fabricated in this study were 6.5-inches wide, 12-inches deep and 18-

feet long. All specimens were prestressed with two 0.6-inch diameter 270-ksi low-relaxation strands placed 2” away from the bottom of the section. In addition to the prestressing steel, two #6 Grade 60 deformed reinforcing bars were placed at the top of

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the section. Figure 2-12 depicts the cross sectional details of the specimens (Ruiz et al., 2007).

Figure 2-12 Beam Cross Section Details (adopted from Ruiz et al., 2007) In order to monitor prestress losses with time, vibrating wire gauges were placed

at prescribed locations along the beam. Gauge readings were taken prior to release, after release and at 3, 5, 7, 14, and 28 days after release. From that point forward readings were taken monthly and will continue to be taken monthly through 2008, 2009 and perhaps longer. The readings from the strain gauges allowed the researchers to document the prestress loss with time. Figure 2-13 displays a plot of the measured prestress losses in all ten beams fabricated in this study.

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Figure 2-13 Measured prestress loss versus beam age (Ruiz et al., 2007) As is seen in Figure 2-13, there is no clear separation between prestress losses in

HSC mixtures and losses in SCC mixtures. Additionally, there is very little change in prestress loss after 28-days. From this data, the researchers concluded that “the preliminary results showed little difference in prestress losses between conventional HSC and high strength SCC mixtures” (Ruiz et al., 2007). In addition, the researchers measured similar amounts of elastic shortening losses between beams with HSC and beams with SCC.

In addition to analyzing data from strain gauges, Ruiz et al. (2007) compared their measured prestress losses to estimated prestress losses calculated using the detailed method of the National Cooperative Highway Research Program (NCHRP) Report 496. The NCHRP Report 496 detailed method is discussed in explicit detail in section 2.5.1 of this report. In comparing the measured prestress losses with the losses calculated by the NCHRP 496 method, the researchers found that prestress losses were overestimated by 20% for the beams cast with HSC and 35% for beams cast with SCC. The researchers reported that “the [NCHRP 496] equations are more sensitive to increases in release strengths than were the measured losses for the beams in this research program” (Ruiz et al., 2007). Ruiz et al. mention that if the beams cast with SCC and beams cast with HSC had the same concrete strength at prestress release, the NCHRP 496 report would have calculated very similar amounts of prestress loss. Because the concrete release strength of the SCC mixtures was much lower than that of the HSC mixtures, the NCHRP 496 procedure estimated a greater percentage of prestress loss for the SCC beams.

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The findings of Ruiz et al. (2007) seem to contradict the findings of Gross et al. (2007). It is important to note that the testing of Ruiz et al. (2007) is still underway (at the time this report was written) and will not be completed until a period of at least one year has elapsed. Furthermore, these two research studies focused on smaller beams that may not necessarily be representative of full-scale member behavior.

2.4.6 Burgueño and Bendert, 2007 During the course of this research project, prestressed box beams were fabricated

with SCC and used in a demonstration bridge project in Jackson, Michigan. Prior to the fabrication of the box beams, full-scale flexure and shear tests were performed to ensure adequate performance.

In order to fabricate these prestressed box beams, three separate SCC mixtures were developed along with a standard normally consolidating concrete (NCC) mixture. The three SCC mixtures were proportioned to achieve minimum compressive strengths of 5000-psi at 1 day and 5,500-psi at 28 days. Upon beam fabrication and testing, both the 1 day and 28 day strength requirements were exceeded. Table 2-8 displays the concrete mixture proportions for all four mixture designs.

Table 2-8 Concrete Mixture Proportions (Burgueño and Bendert, 2007)

Material SCC1 SCC2 SCC3 NCC Type III Cement (lbs) 700 700 700 564

Sand (lbs) 1591 1513 1320 1354 Coarse Aggregate (oven dry) (lbs) 1350 1350 1450 1883

Water (lbs) 256 285 320 151 Air (target) (%) 6 6 6 6

w/c 0.37 0.41 0.46 0.38 Air Entraining (oz/cwt) 1 1 1 1.9

HRWR (oz/cwt) 15 12 10.7 11.3 VMA (oz/cwt) 1 2 6 0

The full-scale box beams used in this research were 52-feet long with a cross

section that was 36-inches wide and 27-inches deep. All prestressing steel in the cross section was 0.6” diameter 270-ksi seven-wire low-relaxation strand. Figure 2-14 includes a drawing of the cross section of the box beams.

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Figure 2-14 Box Beam Cross Section (Burgueño and Bendert, 2007)

In total, 17 beams were fabricated. 8 beams were fabricated with SCC for flexural and shear testing, 6 beams were fabricated with SCC to be used in the bridge, and three substitute beams were produced with NCC in case the performance of the SCC beams was unacceptable. Of the 8 beams fabricated for testing, four were tested in flexure (one beam from each mixture design) and four were tested in shear (also one beam from each mixture design). For the flexural tests, the beams were loaded in four point bending with a span of 50-feet. The two applied loads were spaced 8-feet apart at the center of the beam. For the shear testing, the beams were again loaded in four point bending with loads spaced at 8-feet, but the supports were moved inward to create much smaller shear spans. Schematic diagrams of the flexural and shear test setups are displayed in Figure 2-15.

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Figure 2-15 Full Scale Beam Test Setups (Burgueño and Bendert, 2007) For the flexure tests, “the overall response of all beams was essentially equal”

(Burgueño and Bendert, 2007). In addition, the experimentally determined flexural strengths exceeded design nominal capacities as determined by the simplified approach in the 17th edition of the AASTHO Standard Specifications. The ratio of the experimentally determined flexural strength to the nominal design capacity ranged from 1.1 for the NCC beam to 1.06 for the SCC3 beam. For the shear testing, the response was “very similar and in all cases the critical sections reached a shear capacity greater than that of the design beam” (Burgueño and Bendert, 2007). The ratio of the measured shear strength to nominal design strength according to the 17th edition of the AASHTO Standard Specifications ranged from 1.22 for the SCC1 beam to 1.08 for the SCC3 beam.

Due to the favorable results from the flexure and shear testing, the 6 beams cast with SCC were used in the M-50/US-127 bridge of the Grand River in Jackson, Michigan. The beams were oriented in the bridge as shown in Figure 2-16.

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Figure 2-16 Plan view with beam layout of M-50/US-127 bridge (Burgueño and

Bendert, 2007) The four instrumented beams shown in Figure 2-16 were affixed with

thermocouples to measure temperature through the cross section as well as vibrating wire gauges. These instruments were placed at midspan and a quarter-point section with a back-up set of instruments located 1-foot away from midspan. Five readings from the strain gauges during construction of the beams:

1. After concrete placement 2. After prestress release 3. After cutting of top strands (shown in Figure 2-14) 4. After the erection, and 5. After deck casting

After these five measurements were taken, the instruments were set up for in-service monitoring. In order to accomplish this objective, the instruments were hooked up to an on-site datalogger from which data could be downloaded on a monthly basis. The continuous monitoring of the gauges and thermocouples began December 21, 2005. The recorded strain measurements in the bottom flange of the beams at midspan is displayed in Figure 2-17.

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Figure 2-17 Strain monitoring at beam bottom flange at mid-span section (Burgueño

and Bendert, 2007) As is seen in Figure 2-17, the first five data points from each beam are difficult to

interpret. However, after the continuous monitoring began in December of 2005, the change of strain with time has been consistent among all four beams. As reported by the authors, “Investigation on the reasons behind the initial differences and further analyses on the collected data particularly due to temperature effects, is in progress” (Burgueño and Bendert, 2007). At this point in the research Burgueño and Bendert reached the conclusion that:

even though [the] long-term behavior of the SCC prestressed box beams are still to be fully evaluated, their short-term flexural and shear response determined through the testing program indicates that the SCC beams safely satisfy their prescribed design requirements”(Burgueño and Bendert, 2007).

2.4.7 Naito, Parent, and Brunn, 2006 Naito et. al (2006) presented the results of their research on the strength gain,

creep, shrinkage, stiffness, strand-to-concrete bond quality, and ultimate shear and flexural strength in SCC precast concrete bridge members. In order to accomplish this

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goal, the research was conducted in two phases. The first phase dealt with material characteristics and the second phase focused on full-scale girder performance.

In the first phase of the research, an SCC mixture and a conventional high early strength (HESC) mixture were designed. The mixtures were designed to reach 6800 psi within 24 hours of placement and a 28 day compressive strength of 8000 psi. Hardened properties of these concretes were monitored throughout the research program. The compressive strength of the two concretes fell within expected values and both were comparable to each other. Additionally, the splitting tension and modulus of rupture as determined by ASTM standard tests were found to be consistent between the two mixtures. The measured elastic modulus for SCC was found to be lower than that for the conventional concrete mixture and the values calculated using ACI 318 provisions. Creep and shrinkage tests were performed on match-cast 6 x 12-inch cylinders. The SCC mixture exhibited greater shrinkage and creep strains than the conventional mixture. After evaluating both mixtures over the course of the research program, it was concluded that, “early strength-gain properties of the SCC were comparable to those of traditional HESC” (Naito et al., 2006). Additionally, it was reported that the tensile strength was conservatively higher than the estimates obtained by using ACI 318 expressions and that the direct tension capacity and modulus of rupture between the two concretes were comparable.

In the second phase of the research, Naito et al. (2006) investigated the behavior of full-scale bulb-tee girders cast with SCC. The beams were PennDOT standard beams and contained 26, ½-inch 270 ksi low-relaxation strands. A visual representation of the girder cross-section is included in Figure 2-18.

Figure 2-18 Bulb-tee beam cross section (Naito et al., 2006)

In total, four 35-foot long specimens were cast. Two were cast from SCC and the other two were cast from high early strength concrete. After fabrication, the beams were topped with a conventional 8.5-inch thick deck slab. Then, the girders were examined for modulus of elasticity, creep, shrinkage, and transfer length. The modulus of elasticity was measured in three ways: camber, elastic shortening, and elastic response to applied loads. Elastic shortening was measured at prestress transfer with embedded vibrating wire strain gauges at midspan and conventional foil strain gauges affixed to the prestressing strands prior to casting. In addition, surface strains were measured during

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applied loading and used to calculate the modulus of elasticity. Based on the modulus of elasticity calculation techniques just described, the researchers reported that SCC girders displayed a greater stiffness than the girders fabricated with conventional concrete. This result contradicts the cylinder testing completed by Naito et al. in the first phase of this research project. The authors attributed this contradiction to the fact that the cylinders were moist cured per ASTM standards while the beams were not kept moist after the first 24-hours. In addition to modulus calculations, creep and shrinkage was monitored using the vibrating wire strain gauges embedded in the beam. Naito et al. (2006) reported that the SCC girders experienced less creep and shrinkage than the conventionally cast beams. Furthermore, reduction in prestress force over time due to prestress loss was monitored. The SCC girders exhibited less prestress loss than the conventional high strength beams for the first 75-days. Finally, the transfer length of the beams was calculated using strain gauge measurements before and after release. Both SCC and conventional concrete beams exhibited similar transfer lengths of 15.7-inches and 15.8-inches respectively. Both of these values were below the transfer length estimate obtained through the use of PCI provisions.

To monitor the structural performance at ultimate, the girders were set up to fail in three different ways: compressive-flexural failure, shear-flexural failure, and tensile-flexural failure. Loading points and support points were adjusted to achieve these different failure conditions. All loads were applied with a 5000 kip test machine at a quasi-static loading rate. In all the tests, no significant strand slip was measured, and all girders exceeded their estimated capacities. Ultimately, the researchers concluded that, “the studied SCC provides mechanical characteristics that outperform current industry recommendations” (Naito et al., 2006). In addition, the researchers reported that SCC was a viable material for use in prestressed concrete applications. It is important to note that these conclusions were based on the mixture studied in their research and that the behavior of other SCC mixture designs may need separate investigations.

2.5 METHODS TO ANALYZE PRESTRESS LOSS In the current research project (TxDOT Project 5197), forty five TxDOT Type-C

Beams as well as ten TxDOT 4B28 box beams were tested in flexure to experimentally determine their cracking load. Information on the test specimens is presented in Chapter 3 and the details of the experimental test setup are provided in Chapter 4. In order to compare the observed cracking loads from all fifty-five tests, methods to analytically calculate cracking loads were needed. In order to accurately predict the flexural cracking load for a beam specimen without any bias, two methods were used:

1. The Detailed Prestress Loss Method (Tadros et al., 2003) included in National Cooperative Highway Research Program (NCHRP) Report 496

2. AASHTO-LRFD Refined Loss of Prestress Estimate – 2008 Interim Edition (AASHTO, 2008)

It is important to note that the PCI Design Handbook Loss of Prestress Estimate (PCI, 2004) equations were not used to calculate predicted cracking loads. The PCI equations developed by Zia et al. (1979) do not include time as a variable. Therefore, the

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PCI equations would provide prestress loss values representative of the total prestress loss that the specimen will experience in its design life. Since the flexural testing of all beam specimens took place within the first two months of casting, the PCI equations would not provide meaningful estimates of test specimen cracking loads.

Both the Detailed Prestress Loss Method (Tadros et al., 2003) and the AASHTO-LRFD Refined Loss of Prestress Estimate (AASHTO, 2008) include refined equations to account for instantaneous elastic shortening loss as well as time dependent losses due to creep and shrinkage of concrete and relaxation of prestressing strands. Figure 2-19 displays the typical loss and gain of prestress over time for a bridge girder. For the purposes of this research, only the elastic shortening loss during prestress transfer and the additional loss due to creep, shrinkage and relaxation before deck placement (at the time of flexure test) is of interest.

Figure 2-19 Stress in strands vs. time for a Prestressed Girder (Tadros et al., 2003)

Prior to the adoption of the NCHRP 496 procedure into the AASHTO-LRFD specification, some conservative, simplifying assumptions were made to aid the designer. This results in two similar, but somewhat different, procedures. The details of these two procedures to estimate of prestress loss are discussed further in Sections 2.5.1 and 2.5.2. These two procedures will be referred to as the NCHRP 496 method and the AASHTO-LRFD method respectively for the remainder of this document.

2.5.1 NCHRP Report 496 Detailed Prestress Loss Method In 2003, National Cooperative Highway Research Program (NCHRP) published

Report 496: Prestress Losses in Prestensioned High-Strength Concrete Bridge Girders. The report was meant to, “develop design guidelines for estimating prestress losses in high-strength pretensioned concrete girder bridges” (Tadros et al., 2003). Tadros et al. evaluated the material properties and prestress loss data from seven full-scale bridge

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girders from four different states. One of the seven girders was from the Harris County FM 1960 underpass in Texas. Based on this experimental data and other data assembled from a literature search, formulas were developed to accurately evaluate material properties and prestress loss for high-strength concrete bridge girders.

The authors concluded that local material properties significantly affect the losses due to shrinkage and creep, and modulus of elasticity (Tadros et al., 2003). It is for this reason that two correction factors (K factors) were included into the material properties equations for modulus of elasticity, shrinkage and creep (Table 2-9). The first factor, K1, represents the difference between local average material properties and the national average. The second value, K2, corrects the equations to provide an upper bound, lower bound, or average value. Table 2-9 displays the material properties equations developed from the NCHRP report.

Table 2-9 NCHRP Report 496 Material Properties Equations (Tadros et al., 2003)

Property Equation Variables

Modulus of

Elasticity 33,000 0.140 1000 .

=Modulus of Elasticity (ksi) =Correction factor corresponding to local

material variability =Correction factor corresponding to an upper

or lower bound =Concrete strength at time of interest (ksi)

Shrinkage Strain

480 10 61 4 1064 94735 2.00 0.0143 51

=Concrete shrinkage strain and =same values as in Modulus Equation =Time-development factor

=size factor =humidity factor for shrinkage

=concrete strength factor ⁄ =Volume to Surface Area Ratio (in) t=age of concrete after loading (days) H= relative humidity (%)

Creep Coefficient

, 1.90 1.56 0.008 .

, =Concrete creep coefficient and =same values as in Modulus Equation , , =same as above =humidity factor for creep =loading factor

= age of concrete after loading (days) =age of concrete at loading (days)

Refined equations for prestress loss were developed based on the material

properties equations from Table 2-9. Tadros et al. (2003) developed equations for the loss of prestress for the entire life of a bridge girder through prestress transfer, deck placement, and service conditions. Since the girders in this research were tested

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individually without a deck, only the prestress loss equations between prestress transfer and deck placement (time of flexure test) were needed.

The NCHRP 496 equations, given in Table 2-10, represent the state-of-the-art for accurately estimating time-dependent prestress losses. It is important to note that the NCHRP 496 method uses transformed section properties to estimate loss due to elastic shortening. This loss is implicitly accounted for by using transformed section properties and the initial prestressing force just prior to transfer. Calculating elastic shortening loss in this manner is more cumbersome but does not require iteration or an estimate of the prestress force after transfer. The equations for prestress loss recommended by NCHRP Report 496 are listed in Table 2-10.

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Table 2-10 NCHRP Report 496 Equations for Prestress Loss (Tadros et al., 2003)

Prestress Loss Equations Variables

Elastic Shortening

1

= Loss of prestress due to Elastic Shortening = Concrete stress at steel centroid (ksi)

= Modulus of elasticity of prestressing strands (ksi) = Initial modulus of elasticity of concrete (ksi)

= Initial prestressing force just prior to transfer (kip) = Area of the transformed section at transfer (in2)

= Moment of Inertia of transformed section at transfer (in4)

= eccentricity of strands of transformed section at transfer (in)

= Self weight moment (in-kips)

Shrinkage

11 1 ,

1 0.7

= Loss of prestress due to shrinkage = transformed section age-adjusted effective

modulus of elasticity factor = Shrinkage strain from initial loading to time of

flexural test = Modular ratio between prestressing steel and

concrete = tensile reinforcement ratio for initial net section = factor for initial net section properties

= aging coefficient that accounts for concrete stress variability with time , = Ultimate creep coefficient

= Area of the net section (in2) = Moment of inertia of the net section (in4)

= eccentricity of strands of net section (in)

Creep , =Loss of prestress due to creep , , =same values as above , =Creep Coefficient from loading to test (days)

Relaxation

1 3

45 0.55 24 124 1

= Loss of prestress due to relaxation = reduction factor reflecting the decrease in strand

prestressing from creep and shrinkage = Intrinsic Relaxation Loss = stress in strands just after transfer (ksi) = specified yield strength of strands (ksi)

= age of concrete at time of test (days) = age of concrete at prestress transfer (days)

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2.5.2 AASHTO-LRFD Refined Loss of Prestress Estimate – Interim Edition 2008 The American Association of State Highway and Transportation Officials

(AASHTO) publishes the AASHTO-LRFD Bridge Design Specification. This specification is the national standard for bridge design in the United States. The material properties equations in the AASHTO-LRFD (2008) specifications are very similar to those included in NCHRP 496 with a few exceptions. First, the K factors were removed from the shrinkage strain and creep coefficient equations. For the modulus of elasticity calculation, only the K1 factor was adopted. While the absence of K-factors allows for slightly simpler equations, the method becomes less responsive to local materials. The material properties in section 5.4.2 of the 2008 AASHTO-LRFD Bridge Design Specification are presented in Table 2-11.

Table 2-11 AASHTO-LRFD Equations for Material Properties (AASHTO, 2008)

Property Equation Variables

Modulus of

Elasticity

33,000 .

=Modulus of Elasticity (ksi) =Correction factor for source of aggregate =unit weight of concrete (kcf) =Concrete strength at time of interest (ksi)

Shrinkage Strain

0.48 10 1.45 0.13 ⁄ 1.0 2.00 0.014 51

61 4

=Concrete shrinkage strain =Time-development factor

=factor for the effect of V/S ratio =humidity factor for shrinkage

=factor for concrete strength ⁄ =Volume to Surface Area Ratio (in) t=age of concrete after the end of curing (days) H= relative humidity (%)

Creep Coefficient

, 1.9 . 1.56 0.008

, =Concrete creep coefficient , , =same as above =humidity factor for creep

= age of concrete after loading (days) =age of concrete at loading (days)

In addition to providing equations for material properties, the AASHTO-LRFD

(2008) specification includes a refined, time-dependent method to calculate prestress losses. As is listed in the commentary in the AASHTO-LRFD (2008) specification, these equations were based on research by Tadros et al., 2003 (NCHRP Report 496). Therefore, the equations for elastic shortening, creep, shrinkage, and relaxation are similar to those of the NCHRP 496 method. One important difference lies in the calculation of prestress loss due to elastic shortening. To estimate elastic shortening loss, the AASHTO-LRFD (2008) procedure recommends the use of gross section properties with an estimated force of 90-percent of the prestress force before transfer. Using gross section properties is easier for the design engineer but creates a source of inconsistency

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between the NCHRP 496 method and the AASHTO-LRFD (2008) method. Additionally, in the AASHTO-LRFD (2008) procedure, the Kit factor used in the creep and shrinkage loss equations is based on gross section properties. The NCHRP 496 method, however, uses net section properties to calculate Kit. That difference causes a small inconsistency in otherwise identical procedures to calculate losses due to creep and shrinkage. Finally, the AASHTO-LRFD (2008) equation for relaxation loss is a simplified version of the equation presented in the NCHRP 496 method. The AAHSTO-LRFD (2008) specifications in calculating the loss due to steel relaxation provide a simple equation that is easy for designers to use. All the equations for prestress loss in Section 5.9.5 of the AASHTO-LRFD (2008) specification are summarized in Table 2-12

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Table 2-12 AASHTO-LRFD Equations for Prestress Loss (AASHTO 2008)

Prestress Loss Equations Variables

Elastic Shortening

0.9 1

= Loss of prestress due to Elastic Shortening

= Concrete stress at steel centroid (ksi) = Modulus of elasticity of prestressing

strands (ksi) = Modulus of elasticity of concrete at

transfer (ksi) = Initial prestressing force just before transfer

(kip) = Area of the gross section (in2)

= Moment of Inertia of gross section (in4) = eccentricity of strands of gross section (in) = Self weight moment (in-kips)

Shrinkage

11 1 1 0.7 ,

= Loss of prestress due to shrinkage = transformed section coefficient = Shrinkage strain from end of curing to

time of flexural test =Modulus of elasticity of concrete at transfer

(ksi) , = Ultimate creep coefficient = Area of the gross section (in2)

= Moment of inertia of the gross section (in4) = eccentricity of strands of gross section (in)

Creep ,

=Loss of prestress due to creep , , =same values as above , =Creep Coefficient from loading to test (days)

Relaxation 0.55

=30 for low relaxation strands

= Loss of prestress due to relaxation = stress in strands just after transfer (ksi) = specified yield strength of strands (ksi)

2.6 SUMMARY Four main topics were covered in this literature review. First a discussion of the

history and recent research of the allowable compressive stress at prestress release was presented. This serves as a background for understanding the origin of allowable compressive stresses and documents the research (Table 2-2) pertaining to the increase of the current code limits.

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Next, a discussion of the mechanical properties of high strength concrete was analyzed. A thorough discussion on concrete material properties was presented. Additionally, the behavior of high strength concrete subjected to temporary and sustained uniaxial load and methods of quantifying internal damage was documented. An understanding of the response of high strength concrete to axial load provides insight into the behavior of the pre-compressed tensile zone in prestressed concrete girders.

Third, the properties and behavior of self-consolidating concrete (SCC) were presented. Research relating to the hardened properties of SCC as well as the effects of SCC in full-scale bridge beams was presented. Self consolidating concrete has been used in prestressed applications and a thorough understanding of the material is necessary to evaluate the behavior of girders cast with SCC in this research project.

Finally, two methods used to estimate prestress loss in prestressed girders was presented. Both the NCHRP Report 496 detailed prestress loss method and the AASHTO-LRFD (2008) refined loss of prestress method were discussed. These two procedures provide refined, time-dependent estimates of prestress loss and will be used to provide an unbiased and consistent means of estimating the cracking moment.

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CHAPTER 3 Test Specimens

3.1 OVERVIEW For the experimental program in this phase of TxDOT Project 5197, forty five

full-scale TxDOT Type-C beams (40-inch deep I-beams) as well as ten full-scale TxDOT 4B28 box beams (28-inch deep by 48-inch wide box beams) were tested in flexure. All specimens were designed to experience a maximum compressive stress at release of 0.60f’ci to 0.70f’ci. The design and production of the forty five Type C beams and ten 4B28 box beams are discussed in sections 3.2 and 3.3 respectively.

3.2 DESIGN AND FABRICATION OF TXDOT TYPE-C BEAMS In this research study (Phase II of TxDOT Project 5197), forty five TxDOT Type-

C beams as well as 10 4B28 box beams were tested in flexure. This section presents a discussion on the design and fabrication of the forty five Type-C beams. Type-C girders are commonly used in Texas bridges and investigating their behavior provided a meaningful basis for understanding of the impact of early prestress transfer. In an effort to gain comprehensive data of prestressed girders produced throughout the state, three different fabricators produced twelve beams and one fabricator produced nine beams, totaling forty five Type-C beams. For the remainder of this report, the four fabricators will remain anonymous and will be simply listed as fabricator A, B, C and D.

In addition to targeting a range of different beam fabrication practices, two main variables were studied in four beam fabrication plants. First, the coarse aggregate used in the mix design for fabricators A and C was limestone whereas fabricators B and D used hard river gravel. This distinction was made in an effort to determine if the coarse aggregate in the concrete would have a significant effect on the live load performance of the girders. The compressive strength of high strength concrete is heavily dependent upon the strength of the coarse aggregate. Additionally, concrete mixtures using river gravel are typically stiffer than those using limestone (Birrcher and Bayrak, 2007). Varying the coarse aggregate among the beams allowed for the investigation of the role of coarse aggregates in prestress transfer and live load performance. The nominal release strength of the beams was the other primary variable. Beams from fabricator A and B were designed for a nominal release strength of 4000 psi while the beams from fabricator C and D were designed for a nominal release strength of 6000 psi. The beams designed for a nominal release strength of 4000 psi were designated as “Series 1” beams. Likewise, the beams designed for a nominal release strength of 6000 psi were designated as “Series 2” beams. The design release strength was varied in order to reflect TxDOT standard designs while giving a realistic representation of the range of release strengths used in prestressed beam production in Texas. The strand pattern of Series 1 beams and Series 2

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beams are discussed in sections 3.2.1 and 3.2.2 respectively. Table 3-1 summarizes the key features of the 45 beams fabricated in four different precast plants.

Table 3-1 Type-C Beam Fabrication Details

Fabricator Nominal Release Strength Coarse Aggregate Beams Produced A 4000 psi (Series 1) Limestone 12 B 4000 psi (Series 1) Hard River Gravel 9 C 6000 psi (Series 2) Limestone 12 D 6000 psi (Series 2) Hard River Gravel 12

Aside from the strand pattern design, all of the Type-C beams were designed

according to AASHTO LRFD Interim Bridge Design Specifications (2008) as shown in the TxDOT standard Type-C drawings. It is important to note here that in addition to the flexural testing described in Chapter 4, the shear strength of 18 Type-C girders was evaluated. The detailed information on the shear testing is reported by Heckmann and Bayrak (2008). After initial shear testing on a couple of test specimens from fabricator A, a modification was made to the shear reinforcement design of these Type-C beams. It was deemed necessary to reduce the shear reinforcement in order to ensure a shear failure would occur before flexural crushing of the beam. In order to weaken the shear capacity, minimum shear reinforcement was used outside of the bursting regions at the end of the beam. It is important to note that this reduction in shear reinforcement has no impact on the flexural strength or cracking moment of the Type-C beams. It is only presented here for the sake of completeness. This change is discussed in further detail in sections 3.2.3.2, 3.2.3.3, and 3.2.3.4 as well as Heckmann and Bayrak, 2008.

Each full-scale Type-C specimen in this study was 56.5 feet long. Designing and testing these full-scale girders provided results that are not influenced by any potential scaling effects. The cross-sectional dimensions and geometric properties of standard Type-C beams are included in Table 3-2.

Table 3-2 Section Properties of a TxDOT Type-C Beam

Beam Type yt (in) yb (in) A (in2) I (in4) w (lb/ft) C 22.91 17.09 494.9 82,602 516

In order to effectively evaluate the possibility of increasing the current code limit

on compressive stress at release, all Type-C specimens produced were designed such that the maximum compressive stress at prestress transfer was within the range of 0.60f’ci to 0.70f’ci.

3.2.1 Design of Series 1 Beams The TxDOT standard Type C beam design was modified for this research study.

A custom, non-standard strand pattern was used to target a maximum compressive stress

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at release of 0.60f’ci to 0.70f’ci at a nominal concrete strength of 4000 psi. The strand pattern used for Series 1 beams is displayed in Figure 3-1.

Figure 3-1 Strand Pattern: Series 1 Specimens

The strand pattern used in Series 1 beams consisted of twenty-six, ½” diameter, 270-ksi low relaxation strands, four of which were harped. The harped strands pass through two “hold down” points located five feet from the centerline of the beam. A representation of the beam and the harping points is shown in Figure 3-2.

Figure 3-2 Type-C Beam Harped Strands

Fabricators A, B, and C achieved this harping pattern by pulling down on the strands as seen in Figure 3-3. Fabricator D harped the strands by pushing down on them with hollow cylindrical rods as shown in Figure 3-4.

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Figure 3-3 Typical Harping for Fabricator C

Figure 3-4 Typical Harping for Fabricator D (photographs courtesy of David Birrcher)

The arrangement of harped strands shown in Figure 3-2 was carefully selected to produce maximum compressive stresses at the bottom surface of the beam at each of the harping points and at the transfer lengths (60 strand diameters from each end of the beam). Using NCHRP K1 and K2 factors of 1.0 in material properties equations (Table 2-

55

9 in Chapter 2) for beam design, the strand pattern shown in Figure 3-1 was selected to produce a maximum compressive stress of 0.69f’ci when prestress transfer occurred at a concrete compressive strength of 4000 psi. Likewise, the maximum compressive stress was 0.65f’ci when prestress transfer occurred at a concrete strength of 4300 psi. Finally, the maximum compressive stress was 0.60f’ci when prestress transfer occurred at a concrete compressive strength of 4700 psi.

In addition to the transfer length and harping points, compressive stresses at release were calculated at “critical sections”. These critical sections were located 2.5-feet away from the centerline of the beam and represented the end of a 5-foot constant moment region applied to the Type-C beam during flexural testing. This 5-foot constant moment region was also used in the flexural tests by Birrcher and Bayrak (2007) and provided a consistent basis for experimental testing. A complete discussion on the experimental test setup is presented in Chapter 4. The aforementioned critical sections were the locations inside the constant moment region (during the flexure test) that experienced the highest compressive stress at prestress transfer. Table 3-3 displays the calculated stresses at the hold down location, transfer length, and the critical section for Series 1 Beams. Figure 3-5 displays a schematic drawing of the transfer lengths, critical sections, and hold-down points on the Type-C beams.

Table 3-3 Calculated maximum compressive stresses at prestress transfer: Series 1

Beams

Compressive Stress at Release (% f’ci) f’ci (psi) Critical Section Hold Down Point Transfer Length

4000 69.0 69.3 69.2 4300 64.5 64.8 64.6 4700 59.4 59.6 59.4

Figure 3-5 Schematic Drawing of Type-C Beam

As seen in Table 3-3, the Series 1 strand pattern adequately produced a range of compressive stress between 0.60f’ci and 0.70f’ci at a nominal release strength of 4000 psi. Additionally, the tensile stresses did not exceed the tensile stress limit at release ( cif ′5.7 ,with f'ci in psi) anywhere along the beam at any of the targeted release

56

strengths. Therefore, no cracking was observed on the top flange of the beams due to prestress force transfer. A sample shop drawing and release stress calculation of a Series 1 Type-C Beam is included in Appendix A.

3.2.2 Design of Series 2 Beams In order to obtain a maximum compressive stress at release of 0.60f’ci to 0.70f’ci at

a nominal concrete strength of 6000 psi, the standard Type-C beam strand pattern was altered once again. The strand pattern used for Series 2 beams is displayed in Figure 3-6.

Figure 3-6 Strand Pattern: Series 2 Beams

The Series 2 beams consisted of thirty-six, ½” diameter, 270-ksi low relaxation strands, six of which were harped. The harping, or “hold down” points for the Series 2 beams were identical to those in the Series 1 beams and are shown in Figure 3-2. As with the Series 1 beams, the maximum compressive stress at release occurred at the hold down locations and transfer lengths. Using NCHRP K1 and K2 factors of 1.0 in material properties equations for beam design, the strand pattern shown in Figure 3-6 produced a maximum compressive stress at release of 0.70f’ci when prestress transfer occurred at a concrete strength of 5400 psi. Likewise, the maximum compressive stress was 0.65f’ci when prestress transfer occurred at a concrete strength of 5850 psi. Finally, the maximum compressive stress was 0.60f’ci when prestress transfer occurred at a concrete strength of 6400 psi. Table 3-4 displays the calculated stresses at the hold down location, transfer length, and the critical section for Series 2 Beams. All three locations were the same for both Series 1 and Series 2 beams (See Figure 3-5).

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Table 3-4 Calculated maximum compressive stresses at prestress transfer: Series 2 Beams

Compressive Stress at Release (% f’ci) f’ci (psi) Critical Section Hold Down Point Transfer Length

5400 70.2 70.4 70.0 5850 65.1 65.3 65.0 6400 59.9 60.1 59.7

As with the Series 1 beams, the strand pattern used in Series 2 beams provided a

range of compressive stress between 0.60f’ci and 0.70f’ci at a nominal release strength of 6000 psi. Once again, this strand pattern did not result in tensile stresses in excess of

cif ′5.7 (as per TxDOT specifications) anywhere along the beam at any of the target release strengths listed above. Thus, no cracking was observed in the top flanges of the beams upon prestress transfer for Series 2 beams. A sample shop drawing and release stress calculation of a Series 2 Type-C Beam is included in Appendix A.

3.2.3 Fabrication of Type-C Girders In order to generate a wide range of data representative of prestressed beam

fabrication in Texas, four different fabricators produced Type-C beams for this study. Although all fabricators used somewhat different methods and concrete mixtures for beam production, the general process for producing Type-C beams was as follows:

1. Prestressing strands were placed along a beam fabrication line and anchored to a thick steel plate at one end (termed the “Dead End”).

2. Strands were tensioned with hydraulic rams to the appropriate jacking force (within a ±2% tolerance) and anchored to a second steel plate at the opposite end of the beam line (termed the “Live End” see Figure 3-7).

3. All mild reinforcing steel was tied in place, and inspected. 4. Steel forms were placed and secured. 5. Concrete was placed using motorized hopper trucks (sidewinders). 6. Beams were covered with wet burlap or waterproof tarps. 7. At an appropriate concrete strength, prestressing strands were released. The following sections detail the fabrication of all forty five Type-C girders used

in this research study. It is important to note that even though the general process for fabricating a Type-C beam was the same, no two fabricators produced Type-C beams with identical methods. It is for this purpose that the details of beam production for each fabricator are documented in the following sections. Also, a copy of the beam fabrication specifications used by each fabricator is included along with the sample shop drawings in Appendix A.

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3.2.3.1 Production of Girders: Fabricator A Fabricator A produced Series 1 girders using limestone as the coarse aggregate.

The concrete mixture design used to fabricate these beams complied with class H concrete mixture design (TxDOT specifications) and was representative of the standard mixtures used in prestensioned beam production. The components of the mixture design used by fabricator A are listed in Table 3-5. Because the moisture conditions of the fine and coarse aggregate could vary, the exact mixture proportions for each casting were documented on batch tickets. The batch tickets for fabricator A are included in Appendix B.

Table 3-5 Concrete Mixture Design used by Fabricator A (per cubic yard)

Components Mixture Design Source Water / Cement Ratio 0.37 --

Water 207 lbs San Marcos Water Co. Cement 564 lbs Alamo Type III

Fine Aggregate 1486 lbs TXI- Green Pit Coarse Aggregate 1796 lbs Hanson-Ogden (Limestone)

High-range water-reducing admixture 8 oz/100 lbs Sika Viscocrete 2100 Retarding admixture 2.4 oz/100 lbs Sika Plastiment

Theoretical Unit weight 150.11 lbs/ft3 --

All beams produced by fabricator A met AASHTO-LRFD Interim Design Specifications (2008) and TxDOT Standard Type-C Beam details (ftp://ftp.dot.state.tx.us/pub/txdot-info/cmd/cserve/standard/bridge/ibdstde1.pdf) Of the twelve beams produced by fabricator A, three were targeted for a maximum compressive stress at release of 0.70f’ci, six were targeted for 0.65f’ci, and three were targeted for 0.60f’ci. This fabricator produced the twelve beams in two separate castings of three beams (0.60f’ci and 0.70f’ci) and one casting of six beams (0.65f’ci). In order to successfully monitor the concrete strength gain with time, twenty four cylinders were cast with each beam line. These cylinders were prepared immediately after the final beam in the beam line was cast. The cylinders were temperature match-cured. As part of the match-curing system (Sure Cure) thermocouples were placed in the beams on the casting line to match the curing temperature of the cylinders to the beams on the casting line. This process ensured that the cylinders gained strength in the same environment as the beams. Starting at approximately six hours after casting, two cylinders were tested every hour. The average value of these two cylinder tests was taken as the concrete strength. As the concrete strength neared the targeted release strength, pairs of cylinders were tested at shorter time intervals to document the strength gain. Once the compressive strength of concrete within 100-200 psi of the targeted release strength, the beam line was released.

Because class H concrete gains strength quickly within the first twenty-four hours, the prestress release process needed to occur quickly to ensure that the prestress force was transferred to the beams at the targeted release strength. Just prior to release,

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the steel form ties (at the top and bottom of the side forms) were removed and the side formwork was separated from the beams. Large concrete “hold-down” blocks were placed over the beam’s hold down points to avoid any cracking due to the dynamic nature of prestress transfer. When the concrete strength reached its targeted value, two workers flame cut the device holding the strands down at harping points. Immediately after that, the stressing plate at the live end of the beam line was retracted toward the beam specimens, gradually transferring the force from all the strands into the beams. A photograph of the live end of a typical beam line is included in Figure 3-7.

Figure 3-7 Live end of prestressing line (Fabricator A)

For all three castings, the entire release process was completed in approximately fifteen minutes. Immediately after release, two additional cylinders were broken. The average strength of the two cylinders before prestress transfer and the two cylinders after prestress transfer yielded the release strength for the last beam cast (youngest) in the line.

To account for the concrete strength gain between the first and last beam on the line in a given casting, linear interpolation was used. For each beam on the line, the time when concrete first entered the formwork was recorded. This time was designated as the “start” time. Additionally, the time when the last batch of concrete was placed was recorded. This was designated the “finish” time. Therefore, every beam on the line had an individual “start” and “finish” time. Using this time lapse data, the concrete strength at release for each individual beam was calculated through interpolation. Table 3-6 displays the maximum compressive stresses induced on the bottom face of the beams at the critical section for Fabricator A.

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Table 3-6 Beam Production Details: Fabricator A

Beam Mark

Design Release Strength (psi)

Actual Release Strength (psi)

Maximum Compressive Release Stresses2 Date of Cast Targeted

σBOTTOM Actual σBOTTOM

CA-70-1 4000 3940 -0.70*f'ci -0.70*f'ci 9/26/07 CA-70-2 4000 3930 -0.70*f'ci -0.70*f'ci 9/26/07 CA-70-3 4000 3930 -0.70*f'ci -0.70*f'ci 9/26/07 CA-65-1 4300 4370 -0.65*f'ci -0.64*f'ci 10/9/07 CA-65-2 4300 4380 -0.65*f'ci -0.63*f'ci 10/9/07 CA-65-3 4300 4310 -0.65*f'ci -0.64*f'ci 10/9/07 CA-65-4 4300 4340 -0.65*f'ci -0.64*f'ci 10/9/07 CA-65-5 4300 4330 -0.65*f'ci -0.64*f'ci 10/9/07 CA-65-6 4300 4300 -0.65*f'ci -0.65*f'ci 10/9/07 CA-60-1 4700 45401 -0.60*f'ci -0.61*f'ci 10/3/07 CA-60-2 4700 45401 -0.60*f'ci -0.61*f'ci 10/3/07 CA-60-3 4700 45401 -0.60*f'ci -0.61*f'ci 10/3/07

1Start and finish times for beams in this casting were not recorded 2At Critical Section

Upon completion of the prestress transfer, the side forms were moved away from

the beams and the beams were lifted and transported to storage. Type-C beam shipment and storage is discussed in section 3.2.4

3.2.3.2 Production of Girders: Fabricator B Fabricator B produced Series 1 girders using river gravel as the coarse aggregate.

Representative of common concrete mixture designs used in this plant, the class H concrete used by Fabricator B contained supplementary cementitious materials (Type F fly ash). The components of the concrete mixture design used by fabricator B are listed in Table 3-7. The batch tickets reflecting the exact mixture proportions for each casting are included in Appendix B.

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Table 3-7 Concrete Mixture Design used by Fabricator B (per cubic yard)


Water 223 lbs Well Water Cement 658 lbs Alamo Type III 1a

Fine Aggregate 1191 lbs Arena Coarse Aggregate 1789 lbs Eagle Lake

Supplementary Cementitious Material 165 lbs Headwaters, Type F Fly Ash High-range water-reducing admixture 7 oz/100lbs Sika 2100

Retarding admixture 0.5 oz/100lbs Plastiment Theoretical Unit weight 149.11 lbs/ft3 --

The beams were designed in accordance with the AASHTO-LRFD Interim

Specifications (2008) and TxDOT Standard Type-C beam details with one exception. The spacing of the shear reinforcement (R-bars shown in Figure 3-8) was increased to 24 inches outside of the bursting (end) regions of the beam. Figure 3-8 displays this increased stirrup spacing.

Figure 3-8 Altered Shear Reinforcement

The new shear design with increased spacing of the R-bars meets the AASHTO-LRFD (2008) minimum requirements for shear reinforcement. Further information about the testing of these Type-C girders in shear can be found in Heckmann and Bayrak, 2008.

Fabricator B intended to produce twelve beams, but due to unforeseen complications in construction, only nine beams were acceptable to the research program. Table 3-8 displays all the fabrication information for the beam specimens produced by fabricator B.

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Table 3-8 Beam Production Details: Fabricator B

Beam Mark





CB-70-1 5400 4540 -0.70*f'ci -0.62*f'ci 5/6/08 CB-70-2 5400 4360 -0.70*f'ci -0.64*f'ci 5/6/08 CB-70-3 5400 4180 -0.70*f'ci -0.67*f'ci 5/6/08 CB-70-4 5400 4030 -0.70*f'ci -0.69*f'ci 5/6/08 CB-70-5 5400 3880 -0.70*f'ci -0.72*f'ci 5/6/08 CB-70-6 5400 3680 -0.70*f'ci -0.76*f'ci 5/6/08 CB-60-1 5850 4820 -0.60*f'ci -0.59*f'ci 4/23/08 CB-60-2 5850 5010 -0.60*f'ci -0.56*f'ci 4/23/08 CB-60-3 5850 4620 -0.60*f'ci -0.61*f'ci 4/23/08

1At Critical Section (See Table 3-6) As seen in Table 3-8, three beams were targeted for a maximum compressive

stress at release of 0.65f’ci, while the remaining six beams were targeted for a maximum compressive stress at release of 0.70f’ci. These nine beams were fabricated in two separate castings. To monitor the concrete strength gain with time, fabricator B used a temperature match-curing system similar to that of fabricator A. The match-cured cylinders were tested as necessary until the concrete strength was near the targeted value. In an effort to release quickly, the forms had been removed before the concrete reached its targeted value. When it was time to release, the devices used at the hold down points to harp the strands were flame-cut. Then, the stressing block at the live end of the beam line was retracted to release most of the prestressing force in the strands. Finally, the prestressing strands in-between each beam were flame-cut. For both castings, this process was complete within fifteen minutes.

As with fabricator A, the “start” and “finish” time for each beam was recorded and used to calculate the concrete strength at release for each specimen on the beam line. Because this concrete mixture had a water to cement ratio much lower than fabricators A, C, and D, the concrete gained strength with high variability. For the casting on May 6, additional cylinders were taken from the very first beam cast. These cylinders from the first beam were tested just before and just after release of the beam line. The average compressive strength of the cylinders just before release and just after release were used to calculate the strength of the first beam on the line at release. Using the strength from the first beam and the last beam, the release strength for each beam on the beam line was calculated through interpolation. This measure was only used to make absolutely sure that accurate values of release strength were documented for this concrete mixture design that was gaining strength at a rapid pace. The concrete strengths at release for the beams produced by fabricator B are displayed in Table 3-8.

Upon completion of the prestress transfer, the beams were lifted off the beam line and placed in storage on-site. Type-C beam shipment and storage is discussed in section 3.2.4

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3.2.3.3 Production of Girders: Fabricator C Fabricator C produced Series 2 girders using limestone coarse aggregate. The

concrete mixture design used to fabricate these beams was a class H concrete mixture with limestone aggregate representative of standard mixtures used in TxDOT Type-C beam production. The components of the mixture design used by fabricator C are listed in Table 3-9. The exact mixture proportions adjusted for the moisture content at each casting are included with the batch tickets in Appendix B.

Table 3-9 Concrete Mixture Design used by Fabricator C (per cubic yard)


Water 242 lbs BCW Well Cement 658 lbs Capital Type III

Fine Aggregate 1326 lbs River Sand Multisources Coarse Aggregate 1789 lbs Vulcan 1604 Grade 5

High-range water-reducing admixture 7.5 oz/100 lbs Sika 2100 Retarding admixture 1 oz/100 lbs Sika Plastiment

Theoretical Unit weight 148.7 lbs/ft3 -- Fabricator C produced three beams targeted for a maximum compressive stress at

release of 0.70f’ci, six beams at 0.65f’ci, and three beams at 0.60f’ci. For Series 2 beams, this translated into target concrete strengths of 5400 psi, 5850 psi, and 6400 psi respectively. Fabricator C produced the twelve beams in three separate castings of three beams (0.70f’ci), six beams (0.65f’ci), and three beams (0.60f’ci) respectively. To monitor the concrete strength gain, fabricator C made 48 cylinders with each beam line. All cylinders were cured with the beam underneath a waterproof tarp. At the appropriate time, cylinders were removed from underneath the tarp and tested. As with fabricators A, and B, fabricator C tested two cylinders approximately six hours after casting and continued to test cylinders every hour after that. When the concrete strength was within 1000 psi of the target, cylinders were tested at a maximum of every half hour until the target strength was achieved. The results from all three castings by fabricator C are displayed in Table 3-10.

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Table 3-10 Beam Production Details: Fabricator C

Beam Mark





CC-70-1 5400 5380 -0.70*f'ci -0.70*f'ci 3/14/08 CC-70-2 5400 5360 -0.70*f'ci -0.71*f'ci 3/14/08 CC-70-3 5400 5330 -0.70*f'ci -0.70*f'ci 3/14/08 CC-65-1 5850 5970 -0.65*f'ci -0.64*f'ci 3/26/08 CC-65-2 5850 6000 -0.65*f'ci -0.64*f'ci 3/26/08 CC-65-3 5850 6130 -0.65*f'ci -0.62*f'ci 3/26/08 CC-65-4 5850 6070 -0.65*f'ci -0.63*f'ci 3/26/08 CC-65-5 5850 6350 -0.65*f'ci -0.60*f'ci 3/26/08 CC-65-6 5850 6250 -0.65*f'ci -0.61*f'ci 3/26/08 CC-60-1 6400 6350 -0.60*f'ci -0.60*f'ci 4/1/08 CC-60-2 6400 6370 -0.60*f'ci -0.60*f'ci 4/1/08 CC-60-3 6400 6370 -0.60*f'ci -0.60*f'ci 4/1/08

1At Critical Section (See Table 3-6) To prepare for an expedient release of the beam line, the forms were removed

before the concrete strength was near the targeted value in all three castings. Once the concrete strength was within ±100 psi of the target, workers flame-cut the device used at the hold down locations. Immediately after that, the stressing plate at the live end was retracted, transferring the prestressing force into the specimens. Then, the strands between each of the beams were flame-cut to finalize the prestress transfer operation. In all three castings, prestress transfer took place within fifteen minutes. Once the beams were visually inspected, they were lifted off the stressing line and placed in storage until they were shipped to the Ferguson Structural Engineering Laboratory.

3.2.3.4 Production of Girders: Fabricator D Fabricator D produced Series 2 girders using river gravel as the coarse aggregate.

The concrete mixture design used to fabricate these beams was again a class H concrete mixture representative of standard mixtures used in beam production. The components of the mixture design used by fabricator D are listed in Table 3-11. The exact mixture proportions adjusted for the moisture conditions at each casting are included with the batch tickets in Appendix B.

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Table 3-11 Concrete Mixture Design used by Fabricator D (per cubic yard)


Water 196 lbs City of Victoria Cement 611 lbs Alamo Type III

Fine Aggregate 1285 lbs Fordyce Murphy Coarse Aggregate: 1990 lbs Fordyce Murphy

High-range water-reducing admixture 27 oz/100 lbs Rheobuild 1000 Retarding admixture 3 oz/100lbs Pozzolith 300R

Theoretical Unit weight 151.19 lbs/ft3 -- As with fabricators A and C, fabricator D cast three beams targeted at 0.70f’ci, six

beams at 0.65f’ci, and three beams at 0.60f’ci. Fabricator D used four separate castings (3 beams per casting) to achieve these different targeted release strengths. In order to monitor the concrete strength gain with time, twenty four cylinders were prepared after the last beam on the beam line was cast. All twenty four cylinders were temperature match-cured. As with the previously mentioned fabricators, cylinders were tested in a structured manner to transfer the prestressing force at a concrete strength very close to the targeted value. The results from all four castings are shown in Table 3-12.

Table 3-12 Beam Production Details: Fabricator D

Beam Mark





CD-70-1 5400 5580 -0.70*f'ci -0.68*f'ci 3/4/08 CD-70-2 5400 5500 -0.70*f'ci -0.69*f'ci 3/4/08 CD-70-3 5400 5420 -0.70*f'ci -0.70*f'ci 3/4/08 CD-65-1 5850 56701 -0.65*f'ci -0.67*f'ci 3/7/08 CD-65-2 5850 56701 -0.65*f'ci -0.67*f'ci 3/7/08 CD-65-3 5850 56701 -0.65*f'ci -0.67*f'ci 3/7/08 CD-65-4 5850 59401 -0.65*f'ci -0.64*f'ci 3/14/08 CD-65-5 5850 59401 -0.65*f'ci -0.64*f'ci 3/14/08 CD-65-6 5850 59401 -0.65*f'ci -0.64*f'ci 3/14/08 CD-60-1 6400 6320 -0.60*f'ci -0.61*f'ci 3/12/08 CD-60-2 6400 6310 -0.60*f'ci -0.61*f'ci 3/12/08 CD-60-3 6400 6300 -0.60*f'ci -0.61*f'ci 3/12/08

1Strength was not interpolated between beams for this casting 2At Critical Section (See Table 3-6)

For two of the castings from this fabricator, the cylinder tests conducted prior to

release yielded larger values than those from the tests conducted after prestress transfer. For these casting lines, interpolation using these strength values would mean that the first beam cast on the line would have a slightly smaller compressive strength value than the last beam cast on the line. This result is simply unreasonable. It was deemed appropriate

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to assume the strength at release for all beams on the casting line was the average strength of the cylinders tested immediately before and after prestress transfer. This is noted in Table 3-12.

As with the previous fabricators, fabricator D took less than 15 minutes to transfer the prestress force to all beams. After inspection, the beams were lifted and stored on-site until delivery to the Ferguson Structural Engineering Lab.

3.2.4 Shipment and Storage of Type C Girders All Type C girders were stored at their respective fabrication plants until delivery

to the Ferguson Structural Engineering Laboratory. Beams were shipped either individually or in pairs on trucks. Upon arrival, each beam was at least seven days old. In the laboratory, a 25 ton crane and a 50 foot twin-girder lifting beam were used to move the specimens off the truck and into the testing setup. The specimen was picked up at lifting points located 5.75’ from each end. The crane and lifting beam used to move the C-Beam specimens is shown in Figure 3-9.

Figure 3-9 Lifting a Type C Girder at FSEL

3.3 DESIGN AND FABRICATION OF TXDOT BOX BEAMS In addition to the forty five C-beam specimens in this study, ten TxDOT 4B28

box beams were tested in flexure. All ten box beams were produced in two different

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fabrication plants of a single producer designated as fabricator E. All ten beams were designed for a nominal concrete compressive strength of 4100 psi at prestress transfer using four different concrete mixture designs. The purpose of testing ten box beams fabricated with four different concrete mixture designs was to perform proof-testing on a compressive stress limit at release that seemed adequate for all of the C-beam tests presented in this report and other beams previously tested by Birrcher and Bayrak (2007).

In order to obtain comprehensive results, several variables were employed in the fabrication of these beams. First, the coarse aggregate type was varied. Four of the ten box beams were fabricated by using limestone for the coarse aggregate while the remaining six were produced by using hard river gravel. This distinction was made to investigate whether early release was more damaging to limestone concrete mixtures or river gravel concrete mixtures. As well as the coarse aggregate, the type of concrete used to fabricate these beams was varied. Conventional concrete was used to fabricate five beam specimens while five specimens were fabricated by using self-consolidating concrete. There has been significant interest to investigate the properties and behavior of self-consolidating concrete mixtures in prestressed concrete beam production. Varying the type of concrete used in beam production allowed for a comparison between self consolidating concrete and conventional concrete.

In addition to varying the coarse aggregate type and concrete consolidation technique used in fabricating the box beams, various adjustments were made to the standard beam design for a different research objective of investigating shear capacity at skewed ends of box beams (TxDOT Research project 5831). Therefore, all ten beams were fabricated with one square end and one skewed end. Also, the shape of the box beam void was modified in order to investigate the behavior of the concrete end block and its role in transferring shear forces in TxDOT Project 5831. As well as adjusting the void, the shear reinforcement in the beams was reduced to allow for shear failures of the specimens such that shear behavior of the specimens could be studied. Figure 3-10 displays the shape of the box beams and the three different types of voids used in fabrication.

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Figure 3-10 Box Beam Void Geometry

All adjustments made to the beams for shear are outside the scope of this research and are not discussed in this report. Table 3-13 summarizes all the variables in the design and fabrication of these ten box beam specimens.

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Table 3-13 Box Beam Design Details Beam Mark

Target Release Strength

Concrete Type Coarse Aggregate Type

Skewed End Void Geometry

BB-01 4100 psi Conventional Limestone 30° Skew BB-02 4100 psi Conventional Limestone Square BB-03 4100 psi Conventional Hard River Gravel 30° Skew BB-04 4100 psi Conventional Hard River Gravel Square BB-05 4100 psi Conventional Hard River Gravel Half Square/ Half Skew BB-06 4100 psi Self Consolidating Limestone 30° Skew BB-07 4100 psi Self Consolidating Limestone Square BB-08 4100 psi Self Consolidating Hard River Gravel 30° Skew BB-09 4100 psi Self Consolidating Hard River Gravel Square BB-10 4100 psi Self Consolidating Hard River Gravel Half Square/ Half Skew

Each full-scale box beam was designed to have a length on the long end of 40’-

½”. The section properties of a standard TxDOT Box Beam are displayed in Table 3-14.

Table 3-14 Section Properties of a TxDOT 4B28 Box Beam

Beam Type yt (in) yb (in) A (in2) I (in4) w (lb/ft) 4B28 14.38 13.62 678.8 68,745 707*

*Does not include the weight of the concrete end blocks As indicated earlier, to effectively evaluate the possibility of increasing the

current code limit on compressive stress at release, all specimens produced were designed to be released at a nominal strength of 4100 psi such that the maximum compressive stress at release of 0.66f’ci at the critical sections of the beam (See Table 3-16).

3.3.1 Design of TxDOT Box Beams The TxDOT standard box beam strand pattern was altered for this experimental

research. A unique, non-standard strand pattern was used to target a maximum compressive stress at release of 0.66f’ci at a nominal concrete strength of 4100 psi. The strand pattern used for all ten box beams is displayed in Figure 3-11.

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Figure 3-11 Box Beam Centerline Strand Pattern

The strand pattern shown in Figure 3-11 consists of thirty, ½” diameter, 270-ksi low relaxation strands, none of which were harped. Because the box beam geometry does not permit harping strands, strands were debonded along the length of the beam to ensure that the tensile stress induced anywhere along the length of the beam at release was below cif ′5.7 . In order to debond prestressing strands, a plastic tubing was placed over the strand and duct-taped. This tubing served as a bond breaker between the concrete and the prestressing strands. In the beams fabricated for this research, four strands were debonded ten feet from each end and four additional strands were debonded at four feet from each end. Figure 3-12 includes the debonding pattern used in all ten box beams. Figure 3-13 displays a picture of the debonded strands prior to concrete placement.

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Figure 3-12 Debonding pattern for Box Beams (Not to scale)

Figure 3-13 Debonded strands prior to concrete placement

3.3.2 Production of Box Beams: Fabricator E A precaster with two separate fabrication plants was used to make all ten box

beams tested in this research. Five box beams were produced by using conventional concrete and five were fabricated by using self consolidating concrete. Additionally, four

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beams contained limestone coarse aggregate while six beams contained river gravel coarse aggregate. This combination of variables resulted in four different concrete mixture designs. While the conventionally consolidating concrete mixture designs were representative of the concrete used in typical production, mixture designs used in self consolidating concrete were developed by fabricator E for their commercial projects and were never before used in TxDOT beams. The components of all four mixture designs used by fabricator E are included in Table 3-15.

Table 3-15 Concrete Mix Designs used by Fabricator E (per cubic yard)

Components Mix 1 Mix 2 Mix 3 Mix 4 Concrete Type Conventional Conventional SCC SCC

Coarse Aggregate Type Limestone River Gravel Limestone River Gravel Water / Cement Ratio 0.29 0.27 0.355 0.354

Water 229 lbs 222 lbs 240 lbs 246 Type I/II Cement 0 lbs 658 lbs 0 lbs 0 lbs Type III Cement 599 lbs 0 lbs 545 lbs 560 lbs Class F Fly Ash 200 lbs 165 lbs 136 lbs 140 lbs Fine Aggregate 1148 lbs 1033 lbs 1515 lbs* 1498 lbs*

Coarse Aggregate: 1829 lbs 1948 lbs 1530 lbs 1500 lbs High-range water-reducing admixture 3.1 oz/cwt 6 oz/cwt 5.58 oz/cwt 6 oz/cwt

Retarding admixture 3 oz/cwt 1 oz/cwt 1.5 oz/cwt 1.5 oz/cwt Theoretical Unit weight 148.33 lb/ft3 149.11 lb/ft3 146.9 lb/ft3 146.1 lb/ft3

*Fine aggregate blend of river sand and manufactured sand In order to fabricate the ten box beams, fabricator E used four separate castings.

The box beam casting process consisted of the following: 1. Prestressing strands were placed along the beam fabrication line and anchored

to a thick steel plate at one end (termed the “dead” end). 2. Strands were tensioned with hydraulic rams to the appropriate jacking force

(within a 2% tolerance) and anchored to another thick steel plate at the opposite end of the beam line (termed the “live” end).

3. All mild reinforcing for the lower half of the box-beam specimen was placed, tied and inspected.

4. Steel side forms were connected to the soffit form and secured. 5. Concrete was poured in the bottom flange of the box beams using motorized

hopper trucks (or sidewinders). This casting was termed the “stage 1” concrete placement.

6. A Styrofoam void was placed in the beam specimens and the remaining mild reinforcing in the top half of the beam was expediently placed, tied, and inspected.

7. Concrete was placed in the top flanges and webs of the box beams using motorized hopper trucks. This was termed a “stage 2” concrete placement. While stage 2 concrete placement took place, concrete placed in the bottom flange (Stage 1 concrete placement) was still plastic (i.e. initial set did not

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take place) as per TxDOT’s two-stage monolithic concrete placement specifications.

8. Beams were covered with wet burlap or waterproof tarps. 9. At appropriate concrete strength, prestressing strands were released. In order to monitor strength gain with time, 36 cylinders were made from the

stage 1 pour and 12 cylinders were made from the stage 2 pour. All cylinders were made immediately after the last beam on the line was cast. In addition, 24 of the stage 1 cylinders were conditioned by using a temperature match-curing system.

Because this research is focused on the behavior of the pre-compressed tensile zone of pretensioned girders (i.e. the bottom flange), the prestress transfer was based on the compressive strength of the cylinders from the stage 1 concrete placement. Approximately six hours after casting, two match-cured cylinders from the stage 1 pour were tested. Thereafter, match-cured cylinders were tested as needed to target the release strength of 4100 psi. When the stage 1 cylinders were within 500 psi of the target value, two stage 2 cylinders were tested. This was done to understand the concrete strength throughout the beam. Once the stage 1 concrete was within 100-200 psi of the target, the prestressing force was transferred.

The prestress transfer operation was completed by retracting the hydraulic rams attached to the thick steel plate at the live end of the prestressing line. In this way, the prestressing strands were detensioned, and therefore the prestressing force was transferred into the specimens. Subsequently, workers flame cut the prestressing strands between the beams, completing the transfer operation. This process took approximately fifteen minutes for each casting. Immediately after prestress transfer, two stage 1 cylinders and two stage 2 cylinders were broken. The cylinder strengths before and after release were averaged to give the concrete strength at release, f’ci.

In order to account for the concrete strength gain between the first and last beams cast on a line, the start time and finish time for each beam was recorded. Using the time-lapse data between the first and last beam cast on the line, the concrete strength at release for each individual beam was calculated through interpolation. The results from the box beam castings are displayed in Table 3-16. As with Fabricator D, when concrete cylinder tests indicated a slight concrete strength decrease during prestress transfer, interpolation was not used to calculate compressive strength. In this situation, the average of the concrete cylinder strengths before and after prestress transfer was taken as the concrete compressive strength at prestress release for all specimens on the beam line. This is noted in Table 3-16. Immediately after fabrication, all box beams were inspected and stored on-site.

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Table 3-16 Beam Production Details: Fabricator E

Beam Mark

Design Release

Strength (psi)

Actual Release Strength from Stage 1 casting

(psi)

Maximum Compressive Release Stresses2

Actual Release Strength from Stage 2 casting

(psi)

Date of Cast Targeted σBOTTOM

Actual σBOTTOM

BB-01 4100 4140 -0.66*f'ci -0.65*f'ci 3280 8/21/08 BB-02 4100 4070 -0.66*f'ci -0.66*f'ci 3130 8/21/08 BB-03 4100 4120 -0.66*f'ci -0.65*f'ci 23201 7/22/08 BB-04 4100 4220 -0.66*f'ci -0.64*f'ci 23201 7/22/08 BB-05 4100 4170 -0.66*f'ci -0.64*f'ci 23201 7/22/08 BB-06 4100 40601 -0.66*f'ci -0.66*f'ci 4050 8/25/08 BB-07 4100 40601 -0.66*f'ci -0.66*f'ci 3920 8/25/08 BB-08 4100 40401 -0.66*f'ci -0.66*f'ci 32702 8/26/08 BB-09 4100 40401 -0.66*f'ci -0.66*f'ci 32702 8/26/08 BB-10 4100 40401 -0.66*f'ci -0.66*f'ci 32702 8/26/08

1Strength was not interpolated between beams 2Data from cylinders tested after prestress transfer 3At Critical Section

3.3.3 Shipment and Storage of Box Beams All box beams were stored at the fabrication plant until delivery to the Ferguson

Structural Engineering Laboratory. Beams were shipped individually on a truck. Upon arrival, each beam was at least seven days old. In the laboratory, a 25 ton crane and a 50 foot twin-girder lifting beam were used to move the specimens off the truck and into the testing setup. The specimen was picked up at lifting points located 35-feet apart. The crane and lifting beam used to move the box beam specimens is shown in Figure 3-14.

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Figure 3-14 Lifting a Box Beam at FSEL

3.4 SUMMARY Forty-five full-scale Type-C girders and ten full-scale 4B28 box beams were

fabricated and tested in this research study. Five different precast plants were utilized to provide a broad sample of beams (i.e. concrete mixture designs and beam fabrication practices) made throughout the state of Texas.

A non-standard strand pattern was created for the Type-C beams in order to achieve a maximum compressive stress at prestress release between 0.6f’ci and 0.7f’ci. The Type-C girders were split into two categories based on nominal release strength. Twenty one girders were designated as Series 1 beams with a nominal release strength of 4000 psi while the remaining twenty four girders were designated as Series 2 beams with a nominal release strength of 6000 psi. In essence, the primary difference between Series 1 and Series 2 beams lies in the cementitious materials content of the mixture designs. In addition, twenty one girders (9 Series 1 beams + 12 Series 2 beams) were produced with river gravel as the coarse aggregate and twenty four girders (12 Series 1 beams + 12 Series 2 beams) were produced with limestone as the coarse aggregate.

As with the Type-C girders, a non-standard strand pattern was used to target a maximum compressive stress at release of 0.66f’ci at a concrete compressive strength of 4100 psi. Previously, scaled test specimens from TxDOT research project 4086 as well as Type-A beams and Type-C beams had performed well when the maximum compressive stress at prestress transfer was kept below 0.65f’ci. The compressive stress

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at release for the box beams was held constant in order to verify that a potential increase of the stress limit to 0.65f’ci would be acceptable for a variety of concrete mixture designs. The coarse aggregate and concrete type was varied among the ten specimens. Five specimens were produced by using conventional high strength concrete and the remaining five with self consolidating concrete. Additionally, six beams (3 beams with SCC and 3 beams with conventional concrete) contained hard river gravel as the coarse aggregate while four contained limestone as the coarse aggregate (2 beams with SCC and 3 beams with conventional concrete).

All fifty five test specimens were fabricated and tested within sixteen months as part of a modification made to extend TxDOT project 5197. This extension involved a wide-ranging set of concrete mixture designs, and beam fabrication practices used to produce C-Beams and Box-Beams. The fabrication of test specimens was discussed in Chapter 3. The experimental program is discussed in Chapter 4.

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CHAPTER 4 Test Setup

4.1 OVERVIEW The experimental program for part I of phase II of TxDOT project 5197 consisted

of the flexural static testing of the following types of beams: • Full-scale Series 1 Type-C girders (21 tests: 12 beams from Fabricator A

+ 9 beams from Fabricator B) • Full-scale Series 2 Type-C girders (24 tests: 12 beams from Fabricator C

+ 12 beams from Fabricator D) • Full-scale 4B28 Box Beams (10 tests: 10 beams from Fabricator E)

The full-scale static testing of the Type-C girders described in Section 3.2 was very similar to the full-scale testing on Type-A beams within the previous phase of TxDOT project 5197 (Birrcher and Bayrak 2007). The Type-C beam testing consisted of loading the beams up to approximately 30% above the observed cracking load. The load was applied in a way to produce a constant moment region of five feet at the center of the simply supported beam span. During the tests, the applied load was constantly monitored, and deflection measurements were taken at midspan and at both supports.

The full-scale static testing of the 4B28 Box Beams described in Section 3.3 is similar to that of the Type-C testing. The testing consisted of loading the beams up to approximately 20% above the observed cracking load. Like the Type-C beam tests, the load was applied to produce a five foot constant moment region in the center of the beam. The applied load, midspan deflection and support deflections were constantly monitored during the tests.

In addition to the flexural static testing of the Type-C girders 18 beams were failed in shear following the flexural test. The loading protocol and test setup for the shear testing of TxDOT Project 5197 is described in Heckmann and Bayrak, 2008. Likewise, all ten 4B28 box beams are scheduled to be tested in shear (TxDOT Project 5831). The results of the box beam shear tests will be reported at a later date and do not appear in this report and are considered to be beyond the scope of this research.

The flexural static testing of Type-C girders and 4B28 Box beams are discussed in Sections 4.2 and 4.3 respectively.

4.2 TESTING OF FULL-SCALE TYPE-C GIRDERS Static flexural testing of forty five TxDOT Type-C girders was performed to

experimentally evaluate their cracking load. The loading protocol, test setup, and instrumentation and data acquisition used in these tests are described in the following sections.

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4.2.1 Load Protocol All Type-C Girders were loaded up to approximately 30% above their observed

cracking load. The typical load protocol for all Type-C beams was to load steadily with brief pauses to approximately 85% of the predicted cracking load. At this point, loading was suspended to visually inspect the beams for cracking. Then, the beam was loaded by increments of 2 to 5 kips with intermittent pauses to inspect the beam. During testing, all visually observed cracks were marked on the beam and crack widths were measured using a crack comparator card. Once the load exceeded the observed cracking load by approximately 10%, loading proceeded by increments of 10 to 20-kips. Within the last load stage, cracks were marked and crack widths were measured. After this point, the beam was completely unloaded. Pictures were taken during all loading breaks to document the crack propagation during testing.

For typical static tests of Series 1 beams, loading began steadily with brief pauses up to 90-kips. At 90-kips the beam was visually inspected for cracks. After 90 kips, the load was increased to 100-kips and the beam was again visually inspected. After this load level, the beam was loaded in small increments of 2 to 5 kips until cracking was observed. Once the loading reached 120-kips, loading progressed by increments of 10-kips up to a maximum load of 150-kips. Figure 4-1 presents a visual diagram for the loading program of Series 1 beams.

Figure 4-1 Series 1 Beam Loading Protocol

For typical static tests of Series 2 beams, loading began with brief pauses up to 130-kips. At this point, the beam was visually inspected for cracks. After this, loading progressed in smaller increments of 2 to 5-kips until cracking was observed. Once the loading reached 150-kips, the beam was inspected and then loaded to 160-kips. After

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inspection at 160-kips, the beam was loaded to its maximum value of 180-kips. At this point the beam was thoroughly inspected and all cracks were marked and measured. Figure 4-2 presents a visual diagram of the loading program for Series 2 beams.

Figure 4-2 Series 2 Beam Loading Protocol

4.2.2 Test Setup All forty-five Type-C beams were subjected to four-point loading. Figure 4-3 and

Figure 4-4 display a schematic figure and a photograph of the Type-C beam test setup respectively.

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Figure 4-3 Type-C Beam Test Setup (Not to Scale)

Figure 4-4 Picture of a Type-C Beam Flexural Test Setup

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A reaction frame of two steel columns and two coped wide flange beams was bolted to the strong floor at the midspan of the beam. Two inch thick steel plates were bolted across the top and bottom flanges of the coped wide flange beams to connect them together and provide stability. A single acting hydraulic ram was bolted to the steel plate below the coped wide flange beams. Below the ram, the following items were stacked on top of the beam:

• 400-kip capacity load cell • 8-inch diameter steel spherical head and accompanying steel seat • A combination of 2-inch, 1-inch or ¾-inch steel shim plates • 64-inch long steel spreader beam • Two 3 x14-inch Neoprene Bearing Pads

The neoprene bearing pads sat directly on the top surface of the beam and were placed 30 inches away from the beam centerline. The steel spreader beam was centered on the neoprene pads. Depending on the camber of the test specimen, either a 2” plate or a 1” plate was placed in the center of the spreader beam. The 8”diameter spherical head was placed on top of the steel plate and accounted for any slight eccentricities or non-level surfaces in the test setup. The load cell was placed directly on top of the spherical head followed by either a ¾” plate or a 1” plate; again, depending on the camber of the specimen. Great care was taken in aligning all of these pieces to ensure that the specimen was not loaded eccentrically. To apply load to the specimen, hydraulic fluid was transferred to the ram by a pneumatic pump. As the hydraulic fluid entered the ram, the piston pressed down onto the assembly below, loading the specimen. Figure 4-5 displays a close view of the constant moment region of a Type-C beam test.

Figure 4-5 Midspan Region of a Type-C Beam Test

30” 30” C L

Hydraulic Ram

Load Cell

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All Type-C beams were simply supported. At each end, a two inch thick steel rod was placed between two inch thick steel plates. To achieve a pin condition at one end of the beam, the rod was welded to the steel plate. The roller connection was free to displace at all times. To protect against the beam completely rolling off its support, steel tabs were welded on to the plates of the roller connection. It is important to note that the roller never came into contact with the tabs during a flexural test. Figure 4-6 displays these support conditions.

Figure 4-6 Type C-Beam Roller Support and Pin Support, respectively

In each test, the centerline distance of the pin and roller connections was 55-feet. This allowed for a 9-inch overhang at each end of the beam. A schematic figure and a photograph of the Type-C beam test setup is displayed in Figure 4-3 and Figure 4-4 respectively

4.2.3 Instrumentation and Data Acquisition During all Type-C beam flexural tests, the following items were used to monitor

and collect data: • Two 6-inch linear potentiometers • Two 2-inch linear potentiometers • One 400-kip capacity Load Cell • One Pressure Transducer

All of the devices listed above provide readings in voltages. These voltage readings were then converted to meaningful engineering data by using conversion factors. The output from all six devices was monitored and recorded for each test using computer software.

In all tests, beam deflection was measured by linear potentiometers in three places: at each end of the beam, and at midspan. At each end of the beam, one 2-inch linear potentiometer was placed approximately 6-inches away from the support and approximately 1-inch inside the bottom face of the beam. These linear potentiometers at the ends of the beam were used to ensure that the pin and roller supports did not deflect by unexpectedly large amounts and verify that the supports did not settle during a test. At midspan, two 6-inch linear potentiometers were placed on either side of the beam

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approximately 1-inch away from the edge of the bottom face. The average value of these two linear potentiometers was taken as the midspan deflection of the beam. All four linear potentiometers were carefully placed and plumbed before each test. Figure 4-7 displays a photograph of the two linear potentiometers located at midspan and Figure 4-8 displays a photograph of a linear potentiometer located at a support.

Figure 4-7 Linear Potentiometers Located at Midspan

Figure 4-8 Linear Potentiometer Located at a Support (Type-C Beam)

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The load applied during a test was measured using two devices. The load was explicitly measured using a 400-kip capacity load cell and verified using a pressure transducer. The pressure transducer converted the hydraulic pressure in the hydraulic line into an equivalent applied load. The load cell, however, directly measured the applied load through internal strain gauges. It is for this reason that the load cell readings were taken as the actual load applied while the pressure transducer values were used to verify the load cell’s accuracy.

When the beam was loaded in the range of its predicted cracking load, the load was maintained while the beam was visually inspected for cracks. Hand-held flashlights were used in an attempt to find the earliest crack possible. All visually observed cracks were marked on the beam and crack widths were measured using a crack comparator card. The largest crack width was recorded for each loading increment during the test. Figure 4-9 displays the use of crack comparator card to evaluate crack width.

Figure 4-9 Crack Documentation using a Crack Comparator Card

Once the beams reached the maximum load, the beam was visually inspected, and then the load was completely removed from the beam. This concluded a flexural test of a Type-C girder.

4.3 TESTING OF FULL-SCALE 4B28 BOX BEAMS Static flexural testing of ten TxDOT 4B28 box beams was performed to

experimentally evaluate cracking load. The loading protocol, test setup, and instrumentation and data acquisition in these tests are described in the following sections.

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4.3.1 Load Protocol The box beams were initially loaded with brief breaks up to approximately 85-

percent of their predicted cracking load. After visual inspection, the beams were loaded in increments of 5 kips up to visually observed cracking. Then, loading proceeded in increments of 10 kips up to an ultimate load approximately 20% above the observed cracking load (typically 200-220 kips). At each loading break, the beams were thoroughly inspected for cracks. Any observed cracks were marked and measured using a crack comparator card. Several photos were taken at each loading increment to document the crack propagation during the test. Figure 4-10 displays a visual representation of the loading protocol for the box beam tests.

Figure 4-10 Box Beam Loading Protocol

4.3.2 Test Setup The test setup used for the box beams was very similar to the test setup for the

Type-C beams. All box beams were subjected to four point loading using the same reaction frame and hydraulic ram as described in section 4.2.2. The following items were stacked on top of the beam:

• 400-kip capacity load cell • 8-inch Diameter steel spherical head and accompanying steel seat • ¾-inch and 1-inch steel shim plates • Three 64-inch long steel spreader beams oriented to create a 5-foot

constant moment region at the center of the beam.

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• Four 3 x10-inch Neoprene Bearing Pads The four neoprene pads were centered under the webs of the beam and placed on

top of the specimen. Each neoprene pad was 30-inches away from the center line of the specimen, creating a 5-foot constant moment region during the test. The three blue beams were oriented as shown in Figure 4-11 to split the load applied by the ram and transfer it to the four neoprene pads. A ¾-inch thick plate was placed on top of the blue beams. The 8-inch diameter spherical head was placed on top of the steel plate and accounted for any slight eccentricities or non-level surfaces in the test setup. The load cell was placed directly on top of the spherical head followed by a 1-inch steel plate. As with the Type-C beams, these pieces were carefully aligned to ensure symmetric loading. To apply load to the specimen, hydraulic fluid was transferred to the ram by a pneumatic pump. As the hydraulic fluid entered the ram, the piston pressed down onto the assembly below, loading the specimen.

Figure 4-11 Box Beam Loading Apparatus

As with the Type-C beams, the box beams used simply supported conditions. In order to create these end conditions, a 2-inch thick steel rod was sandwiched between 2-inch thick steel plates to create the pin and roller connections. For the pin connection, the rod was welded to the steel plate below, preventing any lateral translation. The roller connection was not welded, and could move freely. Pictures of the support conditions used in the box beam tests are displayed in Figure 4-12.

Hydraulic Ram

Load Cell

Spherical Seat

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Figure 4-12 Box Beam Roller Support and Pin Support, respectively

After initial testing of the box beams, it was found that the soffit at the pinned end of the beam was not level. Therefore, the beam would not bear completely over the 48-inch long steel plate. To mitigate this problem, the 48-inch long steel plate on top of the pin connection was replaced with a 10-inch long steel plate. This allowed for complete bearing over the 10-inch plate. A photograph of the pinned end of the beam with the 10-inch top plate is displayed in Figure 4-13.

Figure 4-13 Pinned connection with 10-inch long top plate

For each box beam test, the beam was positioned such that the skewed end was part of the overhang beyond the roller support. This placement left a span of 36-feet between the well defined boundary conditions (pin and roller) for each box beam flexure test. A schematic drawing and picture of the test setup is displayed in Figure 4-14 and Figure 4-15 respectively

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Figure 4-14 Box Beam Test Setup (Not to Scale)

Figure 4-15 Picture of Box Beam Flexural Test Setup

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4.3.3 Instrumentation and Data Acquisition The instrumentation used to monitor load, deflection, and support settlement

during flexural testing of box beams was nearly identical to that of the Type-C beams. A thorough discussion of this instrumentation and the data acquisition process was presented in section 4.2.3.

Figure 4-16 displays the two 6-inch linear potentiometers located at midspan of the box beam specimen. As with the Type-C beam tests, additional 2-inch linear potentiometers were located at each support. Because the box beams were four feet wide, the 2-inch linear potentiometers were placed at the center of the bottom flange. Figure 4-17 displays a 2-inch linear potentiometer located near a support.

Figure 4-16 Linear Potentiometers Located at Midspan

Figure 4-17 Linear Potentiometer Located at a Support (Box Beam)

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4.4 SUMMARY Part I of Phase II of TxDOT project 5197, reported herein, consisted of the

flexural static testing of 45 Type-C beams and ten Box Beams. In the tests of the Type-C beams, the load was applied to create a 5-foot constant moment region in the middle of the beam. The load was applied up to approximately 30% above the beam’s observed cracking load. During testing, applied load, midspan deflections, and support deflections were constantly monitored. For the tests of the box beams, the load was also applied to create a 5-foot constant moment region within the middle of the beam. The box beams were loaded up to approximately 20% above their observed cracking load. Applied load, midspan deflections and support deflections were monitored during testing.

In both the Type-C beam tests and the 4B28 box beam tests, loading was paused to visually inspect the beams and document crack formation. All visually observed cracks were marked and measured using a crack comparator card. Photographs were taken to document the crack growth during testing. The results of all static flexural testing are presented in Chapter 5.

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CHAPTER 5 Analysis of Experimental Results

5.1 OVERVIEW This chapter includes a presentation of the results and analysis from the full-scale

flexural tests described in Chapter 4. The experimental research conducted under part 1 of phase II of project 5197 is presented in three sections:

• Type-C beam flexural test results and analysis • 4B28 Box Beam flexural test results and analysis • Experimental research on self consolidating concrete

The experimental results and analysis of the 45 Type-C beam flexural tests are presented in section 5.2. In each flexural test, measured cracking loads were compared to predicted cracking loads. The details of the procedure for measuring the cracking loads experimentally (Section 5.2.1) as well as the methods of predicting the cracking loads (Section 5.2.2) are presented. In order to predict the cracking loads, prestress losses were calculated with the NCHRP Report 496 method as well as the AASHTO-LRFD (2008) method as described in Chapter 2. Equations from both procedures used to calculate the predicted cracking load are presented. Finally, the data generated from this experimental research is compared to the prior flexural testing conducted by Birrcher and Bayrak (2007).

The test results and analysis of the 10 4B28 box beams are presented in Section 5.3. As with the Type-C beam specimens, the measured cracking loads were compared to predicted cracking loads for each box beam tested. The same methods (NCHRP 496 and AASHTO-LRFD 2008) were used to calculate prestress losses and predict cracking loads. The data from the box beam tests is then compared to the data from the Type-C beam tests along with the previous flexural testing by Birrcher and Bayrak (2007).

In addition to the results from the flexural tests, a discussion on the use of self consolidating concrete is presented in Section 5.4. Data from the five box beam specimens fabricated with SCC are presented and compared to those from the five box beam specimens fabricated with conventional concrete. The flexural behavior of beams made with these two types of concretes and observations made during beam fabrication and testing are also presented.

5.2 RESULTS OF TYPE-C BEAM TESTS In this section, the methods used to measure and predict cracking loads in the 45

Type-C beam tests are presented. Section 5.2.1 describes the process of determining a specimen’s cracking load and section 5.2.2 describes the two procedures (NCHRP 496 and AASHTO-LRFD 2008) used to calculate predicted cracking loads. For all 45 Type-C specimens, the accuracy of the predicted cracking load was determined based on the

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measured cracking load from the flexural tests. Finally for each Type-C beam, the accuracy of the specimen’s predicted cracking load was plotted against the specimen’s maximum compressive stress at release (measured at the critical section).

In order to identify each Type-C beam test specimen, a series of letters and numbers was used in the following format: LL-NN-N (Where L=letter and N = number). The first letter for each Type-C specimen was a C indicating that the specimen was a Type-C beam. The next letter was either A, B, C or D and represented the fabrication plant where the beams were produced. The next two numbers represented the target release strength for the beam (60, 65 or 70). Finally, the last number was used to identify a specific beam within a given casting. For example, fabricator A cast three beams with a target release strength of 0.70f’ci on September 26, 2007. The three beams in that casting were designated CA-70-1, CA-70-2, and CA-70-3 respectively.

5.2.1 Measured Cracking Loads In order to measure the cracking load of the test specimens, two primary methods

were used: visual observations, and load deflection analysis. All visually observed cracking was documented during loading breaks in flexural testing. At each load break, several research assistants using flashlights inspected the beam for any cracking. Effort was made to ensure a thorough visual inspection in order to catch any small flexural cracks during testing. The load at which the first crack was found was termed the visually observed cracking load.

In addition to visually observed cracking loads, an analysis of the load-deflection plot from the flexural tests was used to measure cracking loads. This analysis, as shown in Figure 5-1 and Figure 5-2, was a simple procedure that involved drawing two straight lines. The first line traced the initial linear-elastic portion of the load-deflection plot. The second straight line was traced along the portion of the load-deflection plot that indicated the significant departure from linearity. The intersection of these two lines provided the point at which the load deflection plot became non-linear. For each specimen, the load at the intersection of these two lines was taken as the measured cracking load.

The visually observed cracking loads were used to verify the measured cracking loads in this analysis. It is important to note that great effort was taken to ensure the most accurate cracking loads were taken from all 45 flexural tests. Figure 5-1 displays a sample load deflection plot of Series 1 specimen CA-60-3. Figure 5-2 includes a sample load deflection plot of Series 2 specimen CC-70-2. Both of these load deflection plots display the applied load measured by the load cell and the average deflection of the two 6-inch linear potentiometers located at midspan.

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Figure 5-1 Applied Load versus Midspan Deflection for Specimen CA-60-3

Figure 5-2 Applied Load versus Midspan Deflection for Specimen CC-70-2

During the Type-C beam tests, visually observed cracking loads were documented. These were used to verify the measured cracking loads obtained from the

0

20

40

60

80

100

120

140

160

0 0.5 1 1.5 2 2.5 3

App

lied

Load

, kip

Midspan Deflection, in

Pcr =111-kips

0

20

40

60

80

100

120

140

160

180

200

0 0.5 1 1.5 2 2.5

App

lied

Load

, kip


Pcr=138-kips

P2

P2

P2

P2

94

load-deflection plots. The first visually observed cracks occurred within the 5-foot constant moment region. The first cracks were located on the bottom flange of the specimen and progressed upward through the beam as the load was increased. All initial cracks were small and slowly progressed to a typical maximum crack width of 0.010-inches for Series 1 beams and 0.007-inches for Series 2 beams. Figure 5-3 and Figure 5-4 display the first visually observed crack for Series 1 Beam CA-60-1 and Series 2 Beam CD-65-5 respectively. In addition, Figure 5-5 displays a sketch of the typical crack patterns at the maximum applied load for Series 1 and Series 2 beams.

Figure 5-3 Visually observed cracking load for Series 1 Beam CA-60-1

P = 115-kips P = 150-kips

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Figure 5-4 Visually observed cracking load for Series 2 Beam CD-65-5

Figure 5-5 Typical crack pattern at maximum applied load for Type-C Beam tests

Measured cracking loads for all 45 Type-C beams were calculated by analyzing the load deflection plots and verified with the visually observed cracking loads. The measured cracking loads, maximum crack widths, and test dates for all Series 1 and Series 2 tests are displayed in Table 5-1 and Table 5-2 respectively.

P = 147-kips P = 180-kips

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Table 5-1 Measured Cracking Loads for Series 1 Type-C Beams Fabricator Beam Mark Measured Cracking

Load (kip) Maximum Crack

Width at Maximum Applied Load (in)

Date of Test

A CA-70-3 107 0.007 11/1/2007 A CA-70-2 110 0.007 11/15/2007 A CA-70-1 108 0.009 11/27/2007 A CA-60-1 111 0.010 11/29/2007 A CA-60-2 112 0.009 12/4/2007 A CA-60-3 111 0.010 12/12/2007 A CA-65-1 110 0.010 12/18/2007 A CA-65-2 107 0.010 1/4/2008 A CA-65-3 113 0.010 1/8/2008 A CA-65-4 110 0.010 1/10/2008 A CA-65-5 111 0.010 1/15/2008 A CA-65-6 109 0.010 1/17/2008 B CB-60-1 130 0.007 6/19/2008 B CB-60-2 126 0.007 6/23/2008 B CB-60-3 120 0.007 6/24/2008 B CB-70-1 118 0.007 6/25/2008 B CB-70-2 121 0.007 6/26/2008 B CB-70-3 118 0.009 6/27/2008 B CB-70-4 110 0.009 7/1/2008 B CB-70-5 118 0.007 7/2/2008 B CB-70-6 112 0.009 7/7/2008

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Table 5-2 Measured Cracking Loads for Series 2 Type-C Beams Fabricator Beam Mark Measured Cracking

Load (kip) Maximum Crack

Width at Maximum Applied Load (in)

Date of Test

C CC-70-1 142 0.007 5/23/2008 C CC-70-2 138 0.005 5/27/2008 C CC-70-3 139 0.007 5/29/2008 C CC-65-1 136 0.007 5/29/2008 C CC-65-2 141 0.007 6/2/2008 C CC-65-3 135 0.007 6/3/2008 C CC-65-4 134 0.007 6/10/2008 C CC-65-5 141 0.005 6/11/2008 C CC-65-6 142 0.005 6/12/2008 C CC-60-1 130 0.007 6/12/2008 C CC-60-2 130 0.007 6/17/2008 C CC-60-3 138 0.007 6/18/2008 D CD-70-1 142 0.007 4/2/2008 D CD-70-2 143 0.007 4/7/2008 D CD-70-3 147 0.005 4/10/2008 D CD-65-1 143 0.005 4/14/2008 D CD-65-2 135 0.005 4/18/2008 D CD-65-3 137 0.007 4/21/2008 D CD-65-4 136 Hairline 4/24/2008 D CD-65-5 148 0.005 4/29/2008 D CD-65-6 143 0.005 5/1/008 D CD-60-1 143 0.005 5/5/2008 D CD-60-2 146 0.007 5/12/2008 D CD-60-3 142 0.005 5/20/2008

5.2.2 Predicted Cracking Loads In order to estimate cracking loads in an unbiased and consistent manner, two

methods were used: the NCHRP Report 496 method and the AASHTO-LRFD (2008) method. The details of these prestress loss prediction methods were discussed in detail in Chapter 2. Both procedures use time-dependant equations to estimate prestress losses. For the purposes of this research, the following four prestress loss components were considered:

• Elastic Shortening • Prestressing Strand Relaxation • Creep of Concrete • Shrinkage of Concrete

All loss components were calculated from the time of prestress transfer to the time of the flexural test. The loss equations given in the NCHRP 496 method and the AASHTO-LRFD (2008) method were used to generate an expected cracking moment for each beam. The simply supported conditions along with the applied loading pattern in

98

the Type-C beam tests (discussed in Chapter 4) produce the shear force and bending moment diagrams shown in Figure 5-6. Based on the moment diagram and the calculated cracking moment, Equation 5-1 was used to calculate the expected cracking load for Type-C beam specimens.

Figure 5-6 Type-C Beam Specimens: Applied Load, Shear Force, and Bending

Moment Diagrams

Ppredicted=4

Lc-5' Mcr Equation 5-1

where, Lc is the simply supported span length (55-feet for Type-C girders) Wapp is the total weight of the items stacked on top of the test

specimen prior to loading. (Wapp = 0.772 kips for Type-C Specimens)

Mcr is the predicted cracking moment accounting for prestress losses and member self weight (ft-k)

For all Type-C beam specimens, the measured load was compared to the

predicted load using Equation 5-2. This comparison provided a consistent means of

99

determining the accuracy of both the NCHRP 496 method and the AASHTO-LRFD (2008) method.

=

Pmeasured-Ppredicted

Pmeasured100 Equation 5-2

where, Pmeasured is the cracking load determined from physical testing (kip) Ppredicted is the cracking load determined from analysis (kip)

Using Equation 5-2, along with the NCHRP 496 prestress loss calculation

procedure or the AASHTO-LRFD (2008) procedure, the accuracy of the predicted cracking loads were calculated in a consistent manner. Sections 5.2.2.1 and 5.2.2.2 discuss the use of the NCHRP 496 and the AASHTO-LRFD (2008) methods respectively.

5.2.2.1 NCHRP Report 496 Method This section provides the equations and procedures used to calculate the cracking

moment according to the NCHRP 496 procedure. The equations provided by the NCHRP 496 procedure were explicitly followed in determining the cracking moment for each of the 45 Type-C beams tested in flexure.

As mentioned in Chapter 2, both the NCHRP 496 and the AASHTO-LRFD (2008) material properties equations contain a K1 factor to recognize the variation of local materials with respect to the national average. In order to accurately reflect the variation in local material properties from each of the five fabricators in this research, the K1 factor was calculated empirically. The following steps were followed to calculate the K1 factor for each of the four fabricators that produced Type-C girders:

1. Calculate the empirical modulus of elasticity using Equation 5-3 2. Back-calculate the apparent modulus of elasticity (Equation 5-4) by

tracing the initial linear-elastic portion of the load deflection plot. Figure 5-7 displays a sample load deflection plot with the estimated linear elastic portion for specimen CB-70-1.

3. Calculate an average empirical modulus and the average apparent modulus for each fabricator.

4. For each fabricator, divide the average apparent modulus by the average empirical modulus to arrive at the K1 value.

Ec= 33,000K1K2 0.140+f'c

1000

1.5

f'c Equation 5-3

where, K1 and K2 are equal to unity (to reflect the national average empirical modulus)

100

∆mid= Pa

24EcI3L2-4a2 Equation 5-4

where, P= applied concentrated load 2-½ feet away from midspan (kips) L= simply supported span length of specimen (in) a= distance from the support to a concentrated load (in) Ec= modulus of elasticity (ksi) I= moment of inertia of the cross section (in4)

Figure 5-7 Estimated initial slope of load deflection plot to calculate the apparent

modulus of elasticity for CB-70-1

Equation 5-3 is the NCHRP 496 equation for the modulus of elasticity and Equation 5-4 is derived from a linear elastic structural analysis of the flexural test setup. Dividing these two values provides a representation of local material properties relative to the national average. The process of calculating a K1 factor was completed for each of the 45 Type-C beams tested in this study. These values are displayed in Table 5-3.

Table 5-3 Calculated K1 values

Fabricator Average K1 Value A 1.0 B 1.1 C 1.0 D 1.0

0

20

40

60

80

100

120

140

160

0 0.5 1 1.5 2 2.5

App

lied

Load

, kip

s


Load-deflection response

∆mid= Pa

24EcI3L2-4a2

PP

P P

a = 300” a = 300”

L = 660”

101

As shown in Table 5-3, the elastic moduli exhibited in the Type-C beam flexural

tests was close to the national empirical moduli as calculated by Equation 5-3. In addition to the K1 factor, the NCHRP 496 procedure uses a K2 factor to represent an upper bound, lower bound, or average value. For purposes of this research, a K2 value of 1.0 representing the average value was deemed appropriate. Once the K1 and K2 factors were incorporated into the material properties equations as shown in Chapter 2, the prestress loss equations of the NCHRP 496 method could be used accurately. In order to calculate the predicted cracking moment of the test specimens, Equation 5-5 through Equation 5-8 were used. First, Equation 5-5 was used to calculate the bottom-fiber stress of the specimen at prestress transfer. It is important to note that transformed section properties are used in Equation 5-5. Using transformed properties with the prestressing force just prior to release implicitly accounts for the loss component due to elastic shortening.

fcbi= Pi

1Ati

+eptiybti

Iti-Mgybti

Iti Equation 5-5

where, = Initial prestressing force immediately before transfer (kips) Ati = Area of the transformed section at prestress transfer (in2) epti= Eccentricty of the strands within the transformed cross

section at prestress transfer (in) ybti= Distance from beam bottom fiber to centroid of the

transformed section at prestress transfer (in) Iti = Moment of inertia of the transformed section at prestress

transfer (in4) Mg = Self weight moment at midspan (in-k) Once the bottom fiber stresses was computed, the next step was to calculate the

reduction in prestressing force due to time dependant prestress losses. Equation 5-6 displays the equation that can be used to calculate the reduction in prestressing force due to shrinkage, creep, and relaxation losses. Elastic shortening is not included in this equation because it is implicitly accounted for in Equation 5-5.

∆P= ∆fpSR+∆fpCR+∆fpR *Aps Equation 5-6

where, ∆ = Reduction of initial prestress force (kips) ∆ = Loss of prestress due to concrete shrinkage (ksi) ∆ = Loss of prestress due to concrete creep (ksi) ∆ = Loss of prestress due to relaxation of strands (ksi) = Area of prestressing strands (in2)

102

Once the force reduction due to prestress loss is known, the change in bottom

fiber stress due to this force reduction can be calculated. This calculation is shown in Equation 5-7.

∆fcb= ∆PAtt

+∆Pettybtt

Itt Equation 5-7

where, ∆fcb= change in bottom fiber stress due to prestress losses (ksi) ∆ = Loss of initial prestress force (kip) Att= Area of the transformed section at the time of the flexural test

(in2) ett= eccentricity of the strands in the transformed section at the

time of the flexural test (in) ybtt= distance from bottom fiber to centroid of transformed section

at the time of flexural test (in) Itt= Moment of inertia of transformed section at time of flexural

test (in4) Once the change in bottom fiber stress is known, the cracking moment can be

calculated. As shown in Equation 5-8 the change in bottom fiber stress, Δfcb, is subtracted from the bottom fiber stress at prestress transfer, fcbi, and added to the modulus of rupture for concrete, fr. The result gives the bottom fiber stress at which cracking will occur. Multiplying this stress by the moment of inertia of the transformed section and dividing by ybtt gives the predicted cracking moment.

Mcr=

Itt

ybttfcbi-∆fcb+fr Equation 5-8

where, Mcr=Predicted cracking moment (in-kip) fr=Tensile strength of concrete taken as 7.5

1000f'c (f’c in psi, fr in ksi)

Itt, ybtt, fcbi, ∆fcb are defined above Upon calculating the predicted cracking moment from Equation 5-8, the predicted

cracking load was computed using Equation 5-1. This procedure was completed for all 45 Type-C beam flexural tests. For each test, the accuracy of the cracking load prediction was computed using Equation 5-2. Finally for each beam, the accuracy measurement was plotted against the beam’s maximum compressive stress at release (measured at the critical section).

As mentioned in Chapter 1, Phase I of project 5197 (Birrcher and Bayrak, 2007) involved the flexural testing of thirty six specimens. In the same process as the Type-C

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specimens, predicted cracking loads were calculated using NCHRP 496 prestress loss equations. In order to make appropriate conclusions about the live load performance of prestressed girders stressed beyond the allowable limit, the data from Birrcher and Bayrak (2007) must be incorporated with the data generated by the Type-C beams. Combining the data from both phases of the project allows for a complete picture of the test results. The accuracy of cracking load calculations for the 45 Type-C beam specimens along with results from Birrcher and Bayrak (2007) are displayed graphically in Figure 5-8. It is important to note that the NCHRP 496 prestress loss equations were used to calculate predicted cracking loads in the data shown in Figure 5-8.

Figure 5-8 Comparison of measured and predicted cracking loads using NCHRP 496

losses It is important to note that two Type-C beam test specimens (circled in Figure

5-8) exhibited significant sweep upon arrival to Ferguson Structural Engineering Lab. During the flexural tests of these two specimens, the East face of the beam specimens cracked much earlier than the West face of the beam. As shown in Figure 5-8 these two beams exhibited a lower than expected cracking load due to premature cracking on the East face. Because of the nature of this premature cracking, the specimens were not considered to be representative of the behavior of typical Type-C beams.

As seen in Figure 5-8, there is a noticeable decline in accuracy as the concrete compressive stress at release increases. This decline in accuracy indicates that the test

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-5.0

0.0

5.0

10.0

15.0

20.0

25.0

0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

Crac

king

Loa

d Pr

edic

tion

Acc

urac

y, %

Concrete Compressive Stress at Release at Critical Section, *f'ci

Lower Bound

Type-C beams with significant sweep

n=81

104

specimens subjected to higher levels of compressive stress at release exhibited flexural cracking earlier than predicted by the NCHRP 496 process. The data from Figure 5-8 was interpreted using the lower-bound line drawn with respect to all the data. In analyzing the test results displayed in Figure 5-8 it was assumed that a 5% error in predicted cracking load accuracy was acceptable and therefore used as the criteria for beams subjected to a maximum compressive stress at release greater than 0.60f’ci. Based on the lower bound line, the data suggests that a noticeable (greater than 5%) decrease in accuracy occurs at a maximum compressive stress at release of approximately 0.65f’ci.

5.2.2.2 AASHTO-LRFD 2008 Method In addition to the NCHRP Report 496 method, the equations in the AASHTO-

LRFD Interim 2008 specification were used to predict the cracking loads of the test specimens. This section details the AASHTO-LRFD (2008) procedure used to calculate the predicted cracking load for all 45 Type-C girders tested in this study.

As mentioned in Chapter 2, the AASHTO-LRFD (2008) method utilizes the same K1 factor discussed in section 5.2.2.1. Unlike the NCHRP 496 method, the AASHTO-LRFD (2008) procedure only uses this K1 factor for its modulus of elasticity equation. The absence of the K1 factor in the shrinkage strain and creep coefficient equations make the AASHTO-LRFD (2008) procedure easier for designers to use, but less responsive to local materials. Therefore the K1 factors shown in Table 5-3 were only incorporated into the AASHTO-LRFD (2008) modulus of elasticity calculation.

In order to determine the predicted cracking moment according to the AASHTO-LRFD (2008) method, an effective prestressing force and gross section properties were used. Gross section properties aid the designer in simplicity, but cannot be used to implicitly account for elastic shortening. Therefore, in the AASHTO-LRFD (2008) method, a separate elastic shortening loss equation (see Chapter 2) was used. Equation 5-9 displays the equation used to calculate the effective prestressing force.

Peff= ∆fpES+∆fpSR+∆fpCR+∆fpR Aps Equation 5-9

where, Peff= Effective prestressing force (kips) ∆fpES= Loss of prestress due to elastic shortening (ksi) ∆fpSR= Loss of prestress due to concrete shrinkage (ksi) ∆fpCR= Loss of prestress due to concrete creep (ksi) ∆fpR= Loss of prestress due to strand relaxation (ksi) Aps= Area of prestressing strands (in2) Once the effective prestressing force was estimated, the predicted cracking

moment was calculated. Equation 5-10, uses gross section properties along with the effective prestressing force to provide an estimated cracking moment.

105

Mcr= Ig

ybg

Peff

Ag+

Peffepybg

Ig-Mgybg

Ig+fr Equation 5-10

where, Ig= moment of inertia of the gross section (in4)

ybg=distance from the bottom fiber to the centroid of the gross section (in)

Peff= Effective prestressing force (kips) Ag= Area of gross section (in2) ep= Eccentricity of prestressing strands (in) Mg= Specimen self weight moment (in-kips) fr=Tensile strength of concrete taken as 7.5

1000f'c (f’c in psi, fr in ksi)

Once the predicted cracking moment was calculated, Equation 5-1 was used to

determine the predicted cracking load of the specimen. After all the cracking loads were experimentally determined, Equation 5-2 was used to measure the accuracy of the AASHTO-LRFD (2008) predictions. This process was completed for all 45 Type-C beam test specimens in this study. As in section 5.2.2.1, the accuracy of the predicted cracking loads were plotted against the specimens’ maximum compressive stress at release (measured at the critical section).

As with the data from section 5.2.2.1, the data from the Type-C beam tests was combined with the results from Birrcher and Bayrak (2007) in order to obtain comprehensive results and analyze the complete set of test data. The results of the 45 Type-C test specimens using the AASHTO-LRFD (2008) prestress loss equations are displayed along with the results from Birrcher and Bayrak (2007) in Figure 5-9.

106

Figure 5-9 Comparison of measured and predicted cracking loads using AASHTO-

LRFD 2008 losses As with the NCHRP 496 plot (Figure 5-8) the two Type-C beams that exhibited

significant sweep are circled in Figure 5-9. Because of the sweep, it was determined that these two specimens did not adequately represent typical Type-C beam behavior in flexure.

To analyze the test results shown in Figure 5-9, the lower bound line was again used. As with the NCHRP 496 plot, a 5% prediction error tolerance was deemed acceptable and used as the criteria to evaluate the test data. Based on this error tolerance, the data suggests that a substantial loss of prediction accuracy occurs when specimens are subjected to a maximum compressive stress at release greater than 0.7f’ci. Comparing this value with that obtained from the NCHRP plot (0.65f’ci) suggests that the AASHTO-LRFD 2008 procedure would justify a higher allowable compressive stress at prestress transfer. It should be noted, however, that the use of gross section properties in the AASHTO-LRFD (2008) prestress loss calculations creates a source of error. Using transformed section properties in the NCHRP 496 procedure was more responsive to individual test specimen geometry. This difference suggests that the NCHRP 496 procedure may be a more appropriate method to determine an acceptable allowable compressive stress at prestress transfer. A recommendation for increasing the maximum allowable compressive stress at prestress transfer is provided in Chapter 6. A table with the testing details for each of the 45 Type-C beams is included in Appendix C.

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0.0

5.0

10.0

15.0

20.0

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0.0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0

Crac

king

Loa

d Pr

edic

tion

Acc

urac

y, %


Lower Bound


n=81

107

5.3 RESULTS OF BOX BEAM TESTS In this section, the methods used to measure and predict cracking loads for the 10

4B28 box beam tests are presented. The purpose of investigating these box beams was to perform proof testing on an allowable limit that seemed appropriate based on the results from the Type-C beam testing. For this reason, four concrete mixture designs were used and all ten beams were targeted for a maximum compressive stress at release of 0.66f’ci. Additional information about the fabrication of these beams was presented in Chapter 3.

Section 5.3.1 includes a description of the procedure used to determine a box beam specimen’s measured cracking load and Section 5.3.2 includes a summary of the two procedures (NCHRP 496 and AASHTO-LRFD 2008) used to calculate predicted cracking loads. As with the Type-C specimens, the accuracy of the predicted cracking load was determined based on the measured cracking load from the box beam flexural tests. Finally, the data from the box beam tests was plotted with the data from the Type-C beam tests and the data from Birrcher and Bayrak (2007).

In order to uniquely identify each box beam, a numbering system was used. This numbering system was in the following format: LL-NN (L=letter and N=number). The first two letters were “BB” standing for box-beam. The following two numbers indicated a specific box beam (01 through 10). The table containing all the detailed fabrication information about each individual box beam (presented in Chapter 3) is reproduced here for convenience.

Table 5-4 Box Beam Design Details Beam Mark

Target Release Strength

Concrete Type Coarse Aggregate Type

Skewed End Void Geometry

BB-01 4100 psi Conventional Limestone 30° Skew BB-02 4100 psi Conventional Limestone Square BB-03 4100 psi Conventional Hard River Gravel 30° Skew BB-04 4100 psi Conventional Hard River Gravel Square BB-05 4100 psi Conventional Hard River Gravel Half Square/ Half Skew BB-06 4100 psi Self Consolidating Limestone 30° Skew BB-07 4100 psi Self Consolidating Limestone Square BB-08 4100 psi Self Consolidating Hard River Gravel 30° Skew BB-09 4100 psi Self Consolidating Hard River Gravel Square BB-10 4100 psi Self Consolidating Hard River Gravel Half Square/ Half Skew

In addition to experimentally measuring the cracking load of the box beam

specimens, camber measurements and end region cracking due to prestress transfer were documented for each box beam. These items were monitored to determine if the four different mixture designs used to fabricate the beams as well as the maximum compressive stress at prestress transfer (0.66f’ci) would have any impact on the overall quality of the beam. The results of the camber measurements indicated that beams fabricated with SCC contained higher cambers than those fabricated with conventional concrete. The results from the camber measurements are discussed further in section 5.4. No major concerns were noticed with the end region cracking and schematic diagrams

108

documenting the observed cracks in the box beam end regions are presented in Appendix D.

5.3.1 Measured Cracking Loads As with the Type-C beam specimens, both visually observed cracking loads and

analysis of load-deflection plots were used to determine the measured cracking load. The visually observed cracking loads were determined during load breaks. As with the Type-C beams, researchers with flashlights inspected the beam at each loading break to identify and mark cracks. The first observed flexural crack during flexural testing was defined as the visually observed cracking load.

In addition to visual observations, load deflection plots were analyzed to identify the cracking load in the same manner as the Type-C beam specimens. Two straight lines were drawn on the load deflection plot to target the cracking load of the specimen. The first line traced the linear elastic portion of the curve and the second line traced the portion of the curve exhibiting a significant departure from the initial tangent. Figure 5-10 displays the load deflection plot of specimen BB-03 fabricated with conventional concrete containing hard river gravel and Figure 5-11 displays the load deflection plot of specimen BB-10 fabricated with self consolidating concrete containing hard river gravel. Both of these plots display the applied load recorded by the load cell and the average deflection measured by the two 6-inch linear potentiometers located at midspan.

Figure 5-10 Applied Load versus Midspan Deflection for Specimen BB-03

(Conventional Concrete)

0

20

40

60

80

100

120

140

160

180

200

220

240

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

App

lied

Load

, kip


Pcr= 195-kips

P2

P2

109

Figure 5-11 Applied Load versus Midspan Deflection for Specimen BB-10 (SCC)

It is interesting to note the differences between Figure 5-10 and Figure 5-11. Beam BB-03 fabricated with conventional concrete exhibited a much higher modulus of elasticity than beam BB-10. Additionally, beam BB-03 exhibited far less deflection at a load of 200-kips when compared with beam BB-10. These and other differences between self consolidating concrete and conventional concrete are discussed further in Section 5.4.

Visually observed cracking loads were recorded during all box beam flexural tests. As with the Type-C beams, these visually observed cracking loads were used to confirm the loads determined from the load-deflection analysis. The first visually observed cracks occurred on the bottom soffit of the beam in the 5-foot long constant moment region. With increasing load, all cracks propagated up through the bottom flange of the beam and into the web. Flexural testing was stopped once the visually observed cracks reached the mid-height of the web. The typical maximum recorded crack width for the box beams was 0.09-inches. Figure 5-12 displays the visually observed cracks for specimen BB-10

0

20

40

60

80

100

120

140

160

180

200

220

240

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 1.1 1.2

App

lied

Load

, kip


Pcr= 163-kips

P2

P2

110

Figure 5-12 Visually observed cracking load for specimen BB-10 As with the Type-C beam specimens, the measured cracking loads for all 10 box

beams were carefully determined by analyzing the load deflection plots and verified with the visually observed cracking loads. Table 5-5 displays the measured cracking loads, maximum crack widths, and test dates for all 10 box beam specimens.

Table 5-5 Measured Cracking Loads for 4B28 Box Beam Specimens

Fabricator Beam Mark Measured Cracking Load (kip)

Maximum Crack Width at Maximum Applied Load (in)

Date of Test

E BB-01 176 0.007 9/18/08 E BB-02 178 0.009 9/18/08 E BB-03 195 0.013 9/4/08 E BB-04 196 0.010 8/19/08 E BB-05 200 0.007 8/28/08 E BB-06 162 0.010 10/2/08 E BB-07 155 0.007 10/7/08 E BB-08 161 0.009 9/24/08 E BB-09 162 0.007 9/25/08 E BB-10 163 0.009 9/30/08

P = 165-kips Photo of the beam soffit in the 5-

foot constant moment region

P = 200-kips Photo of the West side of the beam in the 5-

foot constant moment region

111

5.3.2 Predicted Cracking Loads As with the Type-C beam specimens, predicted cracking loads were determined

based on the NCHRP Report 496 prestress loss equations and the AASHTO-LRFD 2008 prestress loss equations. These two procedures are outlined in sections 5.3.2.1 and 5.3.2.2 respectively. Both methods use time-dependant equations to estimate the amount of prestress loss in the specimens. As with the Type-C beams, all prestress losses were calculated from the time of fabrication to the time of the flexural test. The simply supported conditions along with the applied loading setup were used to produce the shear force and bending moment diagrams as shown in Figure 5-13. From these conditions, Equation 5-11 was developed to evaluate the predicted cracking load.

Figure 5-13 4B28 Box Beams: Applied Load, Shear Force, and Bending Moment

Diagrams

Ppredicted=4

Lc-5' Mcr Equation 5-11

where, Lc is the simply supported span length (36-feet for Box Beam Specimens)

112

Wapp is the total weight of the items stacked on top of the test specimen prior to loading. (Wapp = 1.972 kips for Box Beam Specimens)

Mcr is the predicted cracking moment accounting for prestress losses and member self weight (ft-k)

For all the box beam specimens, the predicted cracking load from Equation 5-11 was compared with the measured cracking loads using Equation 5-2. This accuracy calculation provided a consistent means of evaluating the predicted cracking loads from the NCHRP 496 prestress loss method and the AASHTO-LRFD (2008) prestress loss method. Section 5.3.2.1 and 5.3.2.2 include a discussion on the calculation of predicted cracking loads using NCHRP 496 prestress loss equations and AASHTO-LRFD (2008) prestress loss equations respectively.

5.3.2.1 NCHRP Report 496 Method This section describes the process used to calculate predicted cracking loads for

the 10 box beam specimens using the NCHRP 496 equations for calculating prestress losses.

In order to calculate material properties for the box beams using the NCHRP 496 method, an appropriate K1 factor was needed to reflect the use of local materials with respect to the national average. As with the Type-C beams, the procedure described in section 5.2.2.1 to calculate the K1 factor was used. Equation 5-3 was used to calculate the empirical modulus of elasticity and Equation 5-4 was used to back-calculate the apparent modulus of elasticity from the initial linear-elastic slope of the load deflection plot. Figure 5-14 displays a sample load deflection plot with the estimated linear elastic portion for specimen BB-05.

113

Figure 5-14 Estimated initial slope of load deflection plot to calculate the apparent

modulus of elasticity for BB-05 Because four different concrete mixture designs were used to fabricate the 10 box

beam specimens, four separate K1 factors were calculated. Table 5-6 displays the concrete type, coarse aggregate type, and calculated K1 factor for each of the 10 box beams.

Table 5-6 Calculated K1 values for Box Beam test specimens

Beam Mark K1 Factor Concrete Type Coarse Aggregate Type BB-01 0.9 Conventional Limestone BB-02 BB-03

1.1 Conventional Hard River Gravel BB-04 BB-05 BB-06 0.8 Self

Consolidating Limestone BB-07 BB-08

0.9 Self Consolidating Hard River Gravel BB-09

BB-10

0

25

50

75

100

125

150

175

200

225

250

0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1

App

lied

Load

, kip


Load-deflection response

∆mid= Pa

24EcI3L2-4a2

PP

P P

a = 186” a = 186”

L = 432”

114

Using the calculated K1 factors from Table 5-6, appropriate predicted cracking loads could be calculated. As described in section 5.2.2.1, Equation 5-5 through Equation 5-8 were used to calculated the predicted cracking moment. Using the predicted cracking moment, Equation 5-11 was used to calculate the predicted cracking load. This process was completed for all 10 box beam specimens.

Once predicted and measured cracking loads were established for all 10 box beams, the accuracy of each predicted cracking load was established for each specimen using Equation 5-2. For each box beam, the accuracy of the cracking load prediction was plotted against the specimen’s maximum compressive stress at release (measured at the critical section). The data from the 10 box beam tests is plotted with the data from the 45 Type-C beams along with the data from Birrcher and Bayrak (2007) in Figure 5-15.

Figure 5-15 Comparison of measured and predicted cracking loads using NCHRP 496

losses

As previously mentioned, the data in Figure 5-15 shows a decrease in cracking load prediction accuracy as the compressive stress at release is increased. The two Type-C beams that exhibited significant sweep are circled, and the five box beams fabricated with self consolidating concrete are also circled. In addition, the lower bound line shown previously in Figure 5-8 is reproduced.

Of the 10 box beams tested, a range of results were observed. The five box beams fabricated with conventional concrete exhibited behavior similar to that of the Type-C

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Loa

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tion

Acc

urac

y, %


Lower Bound


4B28 box beams fabricated with SCC

n=91

115

beam specimens. The conventional beam data was very near or above the lower bound line and therefore did not indicate unconservative predictions of cracking loads. The box beams fabricated with SCC, however, exhibited substantial cracking at loads much lower than predicted by the NCHRP 496 method. The five points circled in Figure 5-15 are clearly below the lower bound line and indicate a significant drop in the accuracy of cracking load estimates. Because of this unconservative result, further research is recommended to understand the behavior of SCC and its applications in prestressed concrete. More information about the observed behavior of the SCC beams is discussed in section 5.4.

If the beams fabricated with SCC are disregarded, the data still suggests that a substantial (greater than 5%) decrease in accuracy of cracking load predictions occurs at a maximum compressive stress at release of approximately 0.65f’ci.

5.3.2.2 AASHTO-LRFD 2008 Method As with the Type-C beam specimens, the equations in the AASHTO-LRFD 2008

specification were used to predict box beam cracking loads. This section covers the AASHTO-LRFD (2008) procedure used to calculate the predicted cracking load for the 10 box beams tested in this study.

In order to calculate the modulus of elasticity using the AASHTO-LRFD (2008) specification, the K1 factors shown in Table 5-6 were used. As previously mentioned, the K1 factor was only used in the modulus of elasticity calculation and is not included in the equations for the shrinkage strain and creep coefficient calculations (as described in Chapter 2 and section 5.2.2.2).

As presented in section 5.2.2.2, Equation 5-9 and Equation 5-10 were used to calculate cracking moments. Using this cracking moment, the predicted cracking load was calculated using Equation 5-11. This process was utilized to calculate the predicted cracking load for all 10 box beams. The accuracy of the predicted cracking load was calculated using Equation 5-2. Figure 5-16 displays the results from the box beam tests using AASHTO-LRFD (2008) prestress loss equations as well as the Type-C beam data along with the data from Birrcher and Bayrak (2007).

116

Figure 5-16 Comparison of measured and predicted cracking loads using AASHTO-

LRFD 2008 losses The data shown in Figure 5-16 exhibit lower cracking load prediction accuracy as

the maximum compressive stress at release is increased. As with Figure 5-15, the two Type-C beams that exhibited significant sweep and the five box beams fabricated with SCC are circled. In addition, the lower bound line drawn in Figure 5-9 is reproduced.

As with the results calculated using NCHRP 496 prestress loss equations, the box beam data shown in Figure 5-16 exhibited a range of results. The five box beams fabricated with conventional concrete all lie above the lower bound line and did not display an unconservative error in predicted cracking load accuracy. Therefore, all five box beams fabricated with conventional concrete exhibited cracking at loads equal to or greater than predicted cracking loads. However, all five beams fabricated with SCC exhibited cracking at loads less than their predicted cracking load. This unconservative result warrants further research into the behavior of prestressed concrete beams fabricated using SCC.

It is important to note that for the box beams fabricated with SCC, the cracking load prediction error using AASHTO-LRFD (2008) equations was smaller than the error using NCHRP 496 equations. Because the AASHTO-LRFD (2008) material properties equations for creep and shrinkage do not include the K1 factor representing local material variability, the AASHTO-LRFD (2008) method predicted greater amounts of prestress

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4B28 box beams fabricated with SCC

n=91

117

loss for the SCC box beams than did the NCHRP 496 method. Because the AASHTO-LRFD (2008) method resulted in greater prestress losses, the AASHTO-LRFD (2008) procedure yielded smaller cracking loads than the NCHRP 496 method. This masked the unconservative nature of the cracking loads obtained for SCC beams when using the AASHTO-LRFD (2008) equations in comparison to the NCHRP 496 equations. Nevertheless, the five beams fabricated using SCC exhibited premature flexural cracking in relation to the five beams fabricated with conventional concrete.

Setting the five box beams fabricated with SCC aside, Figure 5-16, indicates that a substantial loss of prediction accuracy (greater than 5%) occurs when specimens are subjected to a maximum compressive stress at release greater than 0.7f’ci. It is important to note that the AASHTO-LRFD (2008) procedure uses the simplification of gross section properties along with the absence of the K1 factor in material properties equations. As such, the NCHRP 496 equations provide a more appropriate measure to determine a new allowable compressive stress limit at prestress transfer. A recommendation for increasing the maximum allowable compressive stress at prestress transfer is provided in Chapter 6. A table displaying the details for each 4B28 box beam flexural test is provided in Appendix C.

5.4 OBSERVATIONS ON SELF CONSOLIDATING CONCRETE In the fabrication of the 10 box beams tested in this research study, four different

concrete mixture designs were utilized in an effort to gather comprehensive data. During the box beam testing, the five box beams fabricated with self consolidating concrete exhibited significantly different behavior than the beams fabricated with conventional concrete. The following sections describe these differences and provide observations on the use of self consolidating concrete in prestressed concrete beams.

5.4.1 Fabrication of Box Beam Girders During the fabrication of the box beams cast using SCC, honeycombs were

observed on specimens BB-06 and BB-07. Figure 5-17 and Figure 5-18 display samples of the honeycombs from specimens BB-06 and BB-07 respectively.

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Figure 5-17 Sample of honeycombing in box beam BB-06

Figure 5-18 Sample of honeycombing in box beam BB-07

It is important to note that during the fabrication of these two specimens, no external vibration was used to assist the consolidation of the concrete. The SCC mixture in both of these beams displayed unsatisfactory flow properties. In order to have an effective self consolidating concrete mixture, the concrete must be able to flow freely into the formwork leaving a smooth surface finish without serious defects. As evidenced by these two beams, SCC concrete may not always provide the smooth high quality finish required for prestressed concrete construction. In order to avoid this problem with the

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other three box beams fabricated using SCC (BB-08, BB-09, and BB-10) fabricator E decided to provide a very small amount of external vibration after the concrete was placed. While no honeycombing was noticed on those three specimens, external vibration of SCC defeats the primary purpose of a self consolidating mixture and ideally should not be necessary.

Since there was no exposed rebar in the honeycombed areas, these beams were repaired with grout according to TxDOT specifications. Figure 5-19 and Figure 5-20 show box beams BB-06 and BB-07 respectively after the grouting repair. Since the repaired regions were outside of the test region (i.e. the constant moment region) it is believed that the data obtained from test specimens BB-06 and BB-07 are reliable and representative.

Figure 5-19 Repaired Honeycombing of box beam BB-06

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Figure 5-20 Repaired Honeycombing of box beam BB-07

5.4.2 Top Flange Cracking at Release As mentioned in Chapter 3, the box beams were cast in two stages. Because the

focus of the research was on the precompressed tensile zone, the concrete from the stage 1 casting was used to target the concrete release strength of 4100 psi. Due to the fact that on average, the stage 2 concrete was placed one hour after the stage 1 concrete, the tensile strength of the stage 2 concrete was lower at prestress transfer. Therefore, flexural tension cracking in the top flange was observed on all 10 box beam specimens. Although flexural tension cracking would not be desirable in prestressed beam fabrication, the observed cracking in the top flanges had no noticeable effect on the flexural performance of the box beams. Nevertheless, differences in the amount of flexural tension cracks in the top flanges were observed between beams fabricated with SCC mixtures and those fabricated with conventional concrete mixtures. Figure 5-21 displays a typical flexural tension crack observed in a beam containing conventional concrete with hard river gravel along with a typical flexural tension crack observed in an SCC beam containing hard river gravel coarse aggregate.

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Figure 5-21 Typical flexural tension cracking in the top flange of test specimens

As shown in Figure 5-21, the SCC beam (BB-08) exhibited significantly longer tension cracks that beam BB-04. In addition, there were several more cracks along the length of box beam BB-08 within the top flange than there were for beam BB-04. The same phenomenon was observed in the concrete mixtures containing limestone. Figure 5-22 displays a typical top flange crack observed in a conventional concrete beam using limestone coarse aggregate along with a typical top flange crack from an SCC beam using limestone coarse aggregate.

Figure 5-22 Typical flexural tension cracking in the top flange of test specimens

As with the mixtures using hard river gravel, the concrete mixtures with limestone exhibited similar behavior in regard to top flange cracking. It is important to note that for the beams fabricated with concrete mixtures using limestone, the observed difference in top flange cracking between conventional concrete mixtures and SCC mixtures was slight. In general, however, the beams fabricated with conventionally consolidating concrete mixtures exhibited less tensile cracking than those fabricated with SCC. This finding suggests that SCC may have a weaker tensile capacity than conventionally

Conventional BB-04 SCC BB-08

Conventional SCC

BB-01 BB-07

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consolidating concrete. Therefore, the values typically used for design (6 . or 7.5 ) may not be adequate for prestressed elements fabricated with SCC.

5.4.3 Modulus of Elasiticty After each box beam flexural test, load deflection plots were generated using the

readings from the 400-kip load cell and the two 6-inch linear potentiometers located at midspan. From these load-deflection plots, an apparent modulus of elasticity could be determined from the initial linear-elastic portion of the plot. Figure 5-23 displays the load deflection plots for all the box beams cast with concrete containing hard river gravel.

Figure 5-23 Load versus midspan deflection for box beams fabricated with concrete

containing hard river gravel course aggregate As shown in Figure 5-23, the beams fabricated using SCC were significantly less

stiff and they displayed an increased amount of deflection than did those fabricated with conventional concrete. Since the moment of inertia of the cross section is the same for each box beam specimen, this difference in stiffness can be attributed directly to a difference in the modulus of elasticity. In short, the beams fabricated with SCC had a lower modulus of elasticity than the conventional concrete beams. This same trend was observed in the box beam specimens fabricated using limestone as the coarse aggregate.

0

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BB-08 (Self Consolidating Concrete)



Conventional Concrete Specimens

SCC Specimens

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Figure 5-24 displays the load deflection plots for all the box beams cast with concrete containing limestone.

Figure 5-24 Load versus midspan deflection for box beams fabricated with concrete

containing limestone coarse aggregate Although less pronounced in Figure 5-24, the beams cast with SCC were more

flexible (i.e. less stiff) and deflected more under comparable loads than the beams that were conventionally consolidated.

The trends observed in Figure 5-23 and Figure 5-24 suggest that SCC mixtures exhibit a smaller modulus of elasticity than do conventional concrete mixtures. This smaller modulus of elasticity leads to a decreased stiffness and increased deflections under live loading. These factors should be taken into account when using SCC. Further research on the behavioral properties of specimens cast with SCC is recommended to ensure that accurate flexural response can be estimated.

5.4.4 Camber Before every flexural test, the camber of the box beam specimen was measured.

In order to measure camber, a tensioned string was held at each end of the box beam specimen. Once the string was taut and level, the distance between the top surface of the

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beam (at midspan) and the string was measured as the camber. Figure 5-25 displays the camber measurement for specimen BB-05.

Figure 5-25 Camber measurement at midspan for box beam BB-05 (West face)

This process was performed on both faces (East and West) of the beam and the average value was taken as the camber for the specimen. Each of these camber measurements was taken when the specimen was approximately 28-35 days old. Figure 5-26 displays the measured cambers in each of the 6 box beams fabricated with concrete containing hard river gravel coarse aggregate. Figure 5-27 displays the measured cambers in the 4 beams fabricated with concrete containing limestone coarse aggregate.

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Figure 5-26 Measured Camber in box beams fabricated with concrete containing hard

river gravel coarse aggregate

Figure 5-27 Measured camber in box beams fabricated with concrete containing

limestone coarse aggregate As evidenced by the measured cambers, the box beams fabricated with self

consolidating concrete had slightly larger cambers. It is important to note that these

0

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BB-03 BB-04 BB-05 BB-08 BB-09 BB-10

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mbe

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differences in camber are small, but so are the absolute camber values. As a percentage, the differences can be considered more pronounced. This phenomenon may be attributed to the lower modulus of elasticity values observed in the load deflection plots. In effect, it can be anticipated that beams cast using SCC will exhibit higher 28-day cambers than beams cast with typical production conventional concrete.

5.4.5 Cracking Load During all box beam flexural tests, several loading breaks were taken to visually

observe the specimen, document cracks and measure crack widths. During these load breaks, several photos were taken to document the crack patterns and propagation of cracks through the beam specimens. As can be inferred from the load deflection plots (Figure 5-23 and Figure 5-24), the beams fabricated with SCC exhibited significant cracking and non-linear behavior at lower loads than did the beams fabricated with conventional concrete. In the flexural tests, this translated to the SCC beams with more severe cracking than conventional beams at a given load. Figure 5-28 displays the visually observed cracks of companion beams BB-04 (conventional) and BB-08 (SCC) at a load of 200-kips.

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Figure 5-28 Visually observed cracks of beam BB-04 and BB-08 respectively at an

applied load of 200-kips As shown in Figure 5-28, the beam BB-08 fabricated with SCC exhibited

substantially more cracking than did beam BB-04 fabricated with conventional concrete at an applied load of 200-kips. This same trend is observed in beams fabricated with concrete containing limestone coarse aggregate. Figure 5-29 displays the visually observed cracks of beam BB-01 (conventional) and BB-06 (SCC) at a load of 180-kips.

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Figure 5-29 Visually observed cracks of beam BB-01 and BB-06 respectively at an

applied load of 180-kips

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As shown in both Figure 5-28 and Figure 5-29, the beams fabricated SCC exhibited higher amounts of cracking than the beams cast with conventional concrete at the same applied load. This observation is in agreement with the findings from the load deflection plots which indicated that the beams fabricated with SCC displayed larger deflections and a lower modulus of elasticity. Ultimately, it is believed that the box beams fabricated with SCC would be less serviceable and may experience more cracking from overloads than the beams cast with conventional concrete.

5.4.6 Summary of Observations on Self Consolidating Concrete Prestressed concrete box beams fabricated with self consolidating concrete

exhibited substantially different behavior than those fabricated with conventional concrete. The SCC mixtures in this study have has shown improper consolidation, increased amounts of top flange cracking at release, substantially lower modulus of elasticity (along with increased deflections under live loading), slightly higher cambers near 28-days, and lower than expected cracking loads. All of these factors present concerns about the implementation of SCC in prestressed concrete applications. Further research is recommended to examine these concerns with SCC concrete mixtures in future research projects.

5.5 SUMMARY In this phase of TxDOT project 5197, 45 Type-C beams and 10 4B28 box beams

were tested in flexure to experimentally determine their cracking load. The purpose of the testing was to determine the most appropriate maximum allowable compressive stress limit at prestress transfer for prestressed concrete beams. The chapter was presented in three main sections.

The results of the 45 Type-C beam tests were summarized in the section 5.2. Predicted cracking loads were calculated based on prestress loss equations from the NCHRP Report 496 and the AASHTO-LRFD (2008) specification. Measured cracking loads were determined by analyzing load-deflection plots and verified with visual observations during the flexural tests. Based on these measured and predicted cracking loads, the accuracy of the predicted cracking loads was evaluated. Finally, each specimen’s predicted cracking load was plotted against its maximum compressive stress at release. This plot was analyzed using a maximum acceptable error tolerance of 5%. Based on this criteria and the lower bound of the data, a value of 0.65f’ci was found to be appropriate for the NCHRP 496 procedure and a value of 0.7f’ci seemed appropriate for the AASHTO-LRFD (2008) procedure.

Next, the results of the 10 box beam tests were presented. In the same manner as the Type-C beam specimens, predicted and measured cracking loads were determined. Using the predicted and measured cracking loads, the predicted cracking load accuracy was computed. Then the data from the box beams was added to the data from the Type-C beam tests and the data from Birrcher and Bayrak (2007). It was observed that the box beams fabricated with conventional concrete fell within the scatter of the Type-C beams

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and the data from Birrcher and Bayrak (2007). However, the five box beams fabricated with self consolidating concrete exhibited lower cracking loads that were predicted unconservatively. Setting the data from the SCC box beams aside, the previous values of 0.65f’ci for the NCHRP 496 method and 0.7f’ci for the AASHTO-LRFD (2008) method were still acceptable limits for beams cast with conventional concrete.

Finally, observations regarding the use of self consolidating concrete were presented in section 5.4. Because the five beams fabricated with SCC behaved very differently than those fabricated with conventional concrete, several observations were made in regards to beam fabrication, top flange cracking at release, load-deflection analysis, camber measurements, and visually observed cracking. All of these observations indicated concern with the use of SCC in prestressed concrete beam construction.

A summary of this phase of TxDOT project 5197 along with the recommendations for the maximum allowable compressive stress at release are presented in Chapter 6.

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CHAPTER 6 Summary, Conclusions, and Recommendations

6.1 SUMMARY OF EXPERIMENTAL PROGRAM TxDOT Project 5197 was funded in order to investigate the feasibility of

increasing the allowable compressive stress limit at prestress transfer. In phase I of this project, thirty-six specimens were tested to evaluate their live load performance (Birrcher and Bayrak, 2007). Based on the results from this testing, further research was recommended to investigate the behavior of different section types fabricated with several concrete mixture designs. The recommendation by Birrcher and Bayrak (2007) formed the basis for the research of Phase II of TxDOT project 5197.

The work in the current study (part 1 of phase II) involved the static flexural testing of 55 girders. 45 girders were TxDOT Type-C beams (40-inch deep I-beams) and 10 girders were TxDOT 4B28 box beams (28-inch deep by 48-inch wide box beams). In order to obtain a representative range of data, four different fabricators from the state of Texas produced the 45 Type-C beams using conventional concrete with either limestone or hard river gravel coarse aggregate. The maximum compressive stress at release for the Type-C beams ranged from 0.56f’ci to 0.76f’ci. The 10 box beams were fabricated to perform proof-testing on a compressive stress limit that seemed adequate for the Type-C beams. For this purpose, all 10 box beams were targeted for a maximum compressive stress at prestress transfer of 0.66f’ci. The actual maximum compressive stress at release for the box beams ranged from 0.64f’ci to 0.66f’ci. In order to gather comprehensive data, the box beams were fabricated with four different concrete mixture designs:

• Conventional Concrete containing hard river gravel coarse aggregate • Conventional Concrete containing limestone coarse aggregate • Self Consolidating Concrete containing hard river gravel coarse aggregate • Self Consolidating Concrerte containing limestone coarse aggregate

With the use of several different concrete mixture designs along with different cross section shapes, a comprehensive set of data was evaluated in this research study.

In each static flexural test, an experimentally measured cracking load was determined using load deflection plots and visual observations. In addition to the measured cracking loads, predicted cracking loads were analytically determined for all 55 specimens using typical design calculations (P/A ± Mc/I) accounting for prestress losses. Two procedures were used to calculate prestress loss: The NCHRP Report 496 Detailed Prestress Loss Method (Tadros et al., 2003) and the AASHTO-LRFD Interim 2008 Refined Loss of Prestress Estimate (AASHTO, 2008). The accuracy of these predicted cracking loads were then compared to the maximum compressive stress at release for each of the 55 girders. The data from the 55 specimens of Phase II was added to the data from the 36 specimens of Phase I to arrive at a complete set of data from 91 beam specimens.

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Because two separate methods were used to calculate prestress losses in the test specimens, the data was represented in two different ways. As mentioned in Chapter 5, the results calculated using the AASHTO-LRFD 2008 procedure indicated that a maximum compressive stress at release of 0.70f’ci would be adequate. However, the results calculated using the NCHRP Report 496 procedure indicated that a maximum compressive stress of 0.70f’ci would lead to premature cracking and a value of 0.65f’ci would be an appropriate maximum limit for design.

To estimate elastic shortening loss, the AASHTO-LRFD (2008) procedure recommends the use of gross section properties along with an estimated force of 90-percent of the prestress force before transfer. This calculation explicitly accounts for the elastic shortening loss that a pretensioned element will undergo during prestress transfer. The NCHRP 496 method, however, involves the use transformed section properties. Calculating transformed section properties requires more computation, but implicitly accounts for prestress loss due elastic shortening. It is important to note that the AASHTO-LRFD Interim 2008 Bridge Design Specifications do not restrict the designer from using transformed section properties. Section C5.9.5.2.3a of the AASHTO-LRFD Interim 2008 Specifications state:

When calculating concrete stresses using transformed section properties, the effects of losses and gains due to elastic deformations are implicitly accounted for and ∆fpES should not be included in the prestressing force applied to the transformed section at transfer (AASHTO, 2008). Therefore, if a designer chose to use transformed section properties in design, he

or she would essentially be reverting back to the NCHRP 496 Method. Thus, developing conclusions and recommendations from the results derived using the NCHRP Report 496 was deemed appropriate and conservative. It is for these reasons that the conclusions and recommendations discussed in section 6.2 are generated from the results based on predicted cracking loads calculated using prestress losses from the NCHRP Report 496 procedure.

6.2 CONCLUSIONS AND RECOMMENDATIONS After evaluating the live load performance of 91 specimens with different section

types fabricated with different concrete mixture designs, several observations were made: 1) Specimens subjected to maximum compressive stresses greater than 0.65f’ci,

exhibited premature cracking. The compressive stress imposed on these specimens at prestress transfer subjected the concrete in the pre-compressed tensile zone into the non-linear, inelastic range. This caused microcracking in the concrete that was not accounted for in prestress losses or standard design calculations of P/A ± Mc/I.

2) For the 91 specimens tested in TxDOT Project 5197, relaxing the allowable compressive stress at prestress transfer to 0.65f’ci is justified. In general, premature cracking was not observed in specimens subjected to maximum compressive stresses of 0.65f’ci or less. It is important to note that two Type-C beam specimens stressed below 0.65f’ci did exhibit premature cracking but also

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contained significant sweep which impacted their flexural behavior. It was determined that those two specimens did not represent typical behavior of Type-C beams.

3) For the five box beam specimens fabricated with self consolidating concrete, substantial premature cracking was observed. Based on the experimental data from these five box beam tests, the use of self consolidating concrete with a maximum compressive stress at release of 0.65f’ci is not recommended. Based on the conclusions listed above, it is recommended that the requirement

concerning the allowable release stress in AASHTO-LRFD (2008) Bridge Design Specifications should be revised to read:

“Stresses in concrete immediately after prestress transfer (before time-dependent prestress losses): (a) Extreme fiber stress in compression except as permitted in (b) shall not exceed 0.65f'ci. (b) Extreme fiber stress in compression at ends of simply supported members shall not exceed 0.70f'ci”

6.3 RECOMMENDATIONS FOR FUTURE WORK In regards to the live load performance of beams fabricated with self

consolidating concrete, additional testing should be performed. Based on the results of the five box beams fabricated at 0.65f’ci, further testing should be completed to determine an appropriate allowable compressive stress limit where satisfactory live load performance is achieved. Furthermore, an investigation to determine appropriate SCC mixture designs for prestressed applications should be completed. SCC mixture designs should be adjusted to ensure satisfactory material properties (workability, modulus of elasticity, tensile strength, etc.).

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135

APPENDIX A Additional Information for TxDOT Project 5197 Beams

Appendix A includes the following for the TxDOT Project 5197 pretensioned beams: • Type-C Beams

o Sample Series 1 Beam Shop Drawing o Sample Series 2 Beam Shop Drawing o Beam Fabrication Specifications o Series 1 Beam Stress Calculations at Prestress Release o Series 2 Beam Stress Calculations at Prestress Release o Prestress Losses / Cracking Load Calculations

NCHRP 496, and AASHTO LRFD Interim 2008 • 4B28 Box Beams

o Sample Shop Drawing o Stress Calculations at Prestress Release o Prestress Losses / Cracking Load Calculations

NCHRP 496, and AASHTO-LRFD Interim 2008

Figure A-1 Sample Shop Drawing for Series 1 Type-C Beam

136

Figure A-2 Sample Shop Drawing for Series 2 Type-C Beam

137

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Project 5197 Extension Beam Fabrication Specifications

1. Each fabricator will schedule a prefabrication meeting before fabricating any

beams.

2. Each fabricator will submit a concrete mix design to be approved by TxDOT before beam fabrication. The approved mix design will be used to fabricate all 12 beams provided from each fabricator.

3. A ±2% prestressing force and elongation tolerance will be required for all fabricated beams.

4. Each fabricator will start by producing three (3) beams at 0.7f’c. This will be followed by six (6) beams at 0.65f’c and three (3) beams at 0.6f’c.

5. Twenty-four (24) cylinders shall be fabricated with each beam. Of the twenty-four (24), six (6) shall be shipped to the University of Texas with each beam. The remaining eighteen (18) cylinders shall be used to target the required release strength.

6. All cylinders shall be sure-cured.

7. All cylinders shall be marked with the corresponding beam mark (i.e. C70-1) and the date of casting.

8. Approximately six (6) hours after the beam is cast, two (2) cylinders shall be tested in compression every hour. As soon as the strength of the cylinders reaches within 1000-psi of the targeted strength at release, two (2) cylinders shall be tested every thirty (30) minutes or as needed to appropriately achieve the required release strength.

9. When the targeted strength is achieved, all prestressing force shall be transferred to the beam within fifteen (15) minutes.

10. After the beam has been released, two (2) cylinders will be tested in compression.

11.Each Fabricator will report the concrete strength immediately before and immediately after prestress transfer. The time at which the concrete strength was determined will also be reported.

12.Lifting Loops on each beam shall be located 5.75 feet from the end of the beam and be no higher than 5 inches from the top face of the beam. Additional lifting loops may be added if deemed necessary by the fabricator.

Figure A-3 Sample Stress Calculations at Prestress Release for beam CA-70-1 (Series 1)

Beam Mark: CA-70-1Series: Series 1

f'ci (psi) 3929 Type C fcgp (ksi) 2.34Eci (ksi) 3572 Ag (in²) 494.9 n 8.12 Number Height (in) Number Height (in)

K1 1.00 Ig (in4) 82602 ES (ksi) 19.04 8 at 2 8 at 2K2 1.00 yb (in) 17.09 fpo (ksi) 183.46 8 at 4 8 at 4

Eps (ksi) 29000 yt (in) 22.91 Po (kips) 730 6 at 6 6 at 6fpi (ksi) 202.5 ecl (in) 11.86

Strand Dia. (in) 0.5 eend (in) 8.78 2 at 12 2 at 32Aps_1 (in²/N) 0.153 wu (k/ft) 0.516 2 at 14 2 at 34

Ntotal 26 L (ft) 56.5C.G. 5.23 C.G. 8.31Eccentricity 11.86 Eccentricity 8.78

x e Mg

ft ksi in ksi ksi in-k ksi ksiEnd 0.0 0.00 1.47 8.78 1.33 1.78 0 0.00 0.00Transfer Length 2.5 0.04 1.47 9.11 1.38 1.84 417.96 0.09 0.12

5.0 0.09 1.47 9.44 1.43 1.91 797.22 0.16 0.2210.0 0.18 1.47 10.11 1.53 2.05 1439.64 0.30 0.40

Hold Down 23.3 0.41 1.47 11.86 1.79 2.40 2393.4015 0.50 0.66Critical 25.8 0.46 1.47 11.86 1.79 2.40 2451.4515 0.51 0.68Midspan 28.3 0.50 1.47 11.86 1.79 2.40 2470.8015 0.51 0.69

ksi % of f'ci ksiEnd -2.80 71.3 0.30 4.83 OKTransfer Length -2.76 70.4 0.25 4.05 OK

-2.74 69.6 0.22 3.44 OK-2.70 68.8 0.17 2.74 OK

Hold Down -2.77 70.5 0.26 4.18 OKCritical -2.76 70.2 0.25 3.92 OKMidspan -2.75 70.1 0.24 3.84 OK

Stress Calculations at Prestress Release - Type C Beams

Steel

Stress Calculations at Release

Section

Concrete

Centerline EndMaterial Properties Section Properties Elastic Shortening Strand Pattern


Section

Stress at Bottom Fiber Stress at Top Fiber

LRFD Check

cc

c ff

KKE '1000

'140.033000

5.1

21 ⎟⎠⎞

⎜⎝⎛ +=

Lx

APo

IeyP bo

IeyP to

IyM bg

IyM tg

cf '*

IyM

IeyP

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IyM

IeyP

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f tgtootop +−=

g

pg

g

p

gicgp I

eMI

eA

Pf −⎟⎟⎠

⎞⎜⎜⎝

⎛+=

219.0

139

Figure A-4 Sample Stress Calculations at Prestress Release for beam CC-70-1 (Series 2)

Beam Mark: CC-70-1Series: Series 2

f'ci (psi) 5380 Type C fcgp (ksi) 3.19Eci (ksi) 4243 Ag (in²) 494.9 n 6.83 Number Height (in) Number Height

K1 1.00 Ig (in4) 82602 ES (ksi) 21.81 8 at 2 8 at 2K2 1.00 yb (in) 17.09 fpo (ksi) 180.69 8 at 4 8 at 4

Eps (ksi) 29000 yt (in) 22.91 Po (kips) 995 8 at 6 8 at 6fpi (ksi) 202.5 ecl (in) 11.09 6 at 8 6 at 8

Strand Dia. (in) 0.5 eend (in) 8.76 2 at 10 2 at 24Aps_1 (in²/N) 0.153 wu (k/ft) 0.516 2 at 12 2 at 26

Ntotal 36 L (ft) 56.5 2 at 14 2 at 28at at

C.G. 6.00 C.G. 8.33Eccentricity 11.09 Eccentricity 8.76

x e Mg

ft ksi in ksi ksi in-k ksi ksiEnd 0.0 0.00 2.01 8.76 1.80 2.42 0 0.00 0.00Transfer Length 2.5 0.04 2.01 9.01 1.85 2.49 417.96 0.09 0.12

5.0 0.09 2.01 9.26 1.91 2.56 797.22 0.16 0.2210.0 0.18 2.01 9.76 2.01 2.69 1439.64 0.30 0.40

Hold Down 23.3 0.41 2.01 11.09 2.28 3.06 2393.40 0.50 0.66Critical 25.8 0.46 2.01 11.09 2.28 3.06 2451.45 0.51 0.68Midspan 28.3 0.50 2.01 11.09 2.28 3.06 2470.80 0.51 0.69

ksi % of f'c ksiEnd -3.81 -70.9 0.41 5.54 OKTransfer Length -3.78 -70.2 0.36 4.90 OK

-3.75 -69.7 0.32 4.41 OK-3.72 -69.2 0.28 3.87 OK

Hold Down -3.80 -70.6 0.39 5.27 OKCritical -3.79 -70.4 0.37 5.05 OKMidspan -3.78 -70.3 0.36 4.98 OK

Stress Calculations at Prestress Release - Type C Beams

Material Properties Section Properties Elastic Shortening Strand Pattern

Concrete

Centerline End

Steel


Section


Section


LRFD Check

cc

c ff

KKE '1000

'140.033000

5.1

21 ⎟⎠⎞

⎜⎝⎛ +=

Lx

APo

IeyP bo

IeyP to

IyM bg

IyM tg

cf '*

IyM

IeyP

AP

f bgboobot −+=

IyM

IeyP

AP

f tgtootop +−=

g

pg

g

p

gicgp I

eMIe

APf −

⎟⎟⎠

⎞⎜⎜⎝

⎛+=

219.0

140

141

Figure A-5 Prestress Loss/Cracking Load Calculations according to the NCHRP 496

procedure for Specimen CC-70-2, page 1 of 3

DateCast: 3/14/08 ti(days) 1Release: 3/15/08 tt(days) 73

Test: 5/27/08

INPUT

Gross Net Transform Gross Net Transform Number Height (in)A (in²) 494.9 492.97 530.70 494.9 491.5 516.2 8 at 2yb (in) 17.09 17.37 16.56 17.09 17.31 16.77 8 at 4I (in4) 82602 83487 88096 82602 82851 85900 8 at 6

As (in2) 0.612 0 0 0.612 0 0 6 at 8

ys (in) 38.75 38.75 38.75 38.75 38.75 38.75 2 at 10Aps_1 (in²/N) 0.153 0 0 0.153 0 0 2 at 12

N 36 36 36 36 36 36 2 at 14Aps (in

2) 5.508 0 0 5.508 0 0 C.G. 6.00ep (in) 11.09 11.37 10.56 11.09 11.31 10.77yp (in) 6.00 6.00 6.00 6.00 6.00 6.00

Ec (ksi) 4234 4234 4234 6458 6458 6458Eps (ksi) 29000 29000 29000 28883 28883 28883

n 6.85 6.85 6.85 4.47 4.47 4.47

K1 1.0 fpi (ksi) 202.5 L (ft) 56.45833K2 1.0 Aps (in2) 5.508 Span(ft) 55

f'ci (ksi) 5.36 fpy (ksi) 243 H (%) 70f'c_t (ksi) 11.1 fpu (ksi) 270 V (in3) 335294.8

SA (in2) 85220.5fr_t (ksi) 0.790 Pi (kips) 1115.4 V/SA (in) 3.9

Mes. fr_t (ksi) 0Eci (ksi) 4234 Δεp= 0.00698276

Ec_t (ksi) 6458 Δεp-Δεp_ES= 0.00622262

w (pcf) 150w (plf) 516

Mg_span (in-k) 2341.4Mg_max (in-k) 2467.2

Material PropertiesEquations Concrete Steel Additional

Section at MidspanTerm

At Release At Static Test Centerline

Loss of Prestress/Cracking Load Calculations - Series 1 BeamMethod: NCHRP Report 496

Section Properties Strand Pattern

28_'*5.7 cr ff =

cc

c ff

KKE '1000

'140.033000

5.1

21 ⎟⎠⎞

⎜⎝⎛ +=

142

OUTPUT

fcgp= 3.22 ksiΔfpES= 22.0 ksi Kit= 0.79

φi= 0.65Li= 1.43

Initial (fpi) 202.5 ksi ΔfpR= 0.7 ksiAfter ES (fpo) 180.46 ksi

Epst= 28883 ksi

esh= 0.000231K1= 1 Kit= 0.79K2= 1 ρn= 0.01

αn= 1.76ni= 6.85

γsh= 0.48 ψult= 1.41ktd= 0.65 χ= 0.7khs= 1.00ks= 0.94 ΔfpSR= 5.3 ksikf= 0.79

Quantitiy At Test At Ultimateγcr= 0.48 0.74ktd= 0.65 1.00 Kit= 0.79kla= 1.00 1.00 ρn= 0.01ks= 0.94 0.94 αn= 1.76

khc= 1.00 1.00 ni= 6.85kf= 0.79 0.79 ψult= 1.41

K1= 1 1 ψt= 0.91K2= 1 1 χ= 0.7ψt= 0.91 1.41

ΔfpCR= 15.9 ksi

ΔfpES= 22.0 ksiΔfpSR= 5.3 ksiΔfpCR= 15.9 ksiΔfpR= 0.7 ksi

Total ΔfpT= 44.0 ksiTotal-ES ΔfpT-ΔfpES= 21.9 ksi

ShrinkageCreepRelaxation

Elastic Shortening (ES)

Prestressing Force

Shrinkage:Equations Strain Equations Stress

Creep:Equations Equations Stress

Total Losses:

Elastic Shortening: Relaxation:Equations Summary Equations Stress

2190.1),( KKtt cri γψ =

fhcslatdcr kkkkk=γ

118.0−= ila tk

Hkhc 008.056.1 −=

tftk

citd +−

='461

( )735

941064 SV

ks

−=

cif f

k'1

5+

=

ci

psi

n

pnnn

n

psn

ultnniit

ittcgpipCR

EE

n

IeA

AA

nK

Kfnf

=

⎟⎟⎠

⎞⎜⎜⎝

⎛+=

=

++=

=Δ

2

1

)1(11

α

ρ

χψαρ

ψ

itpsshpSR KEf ε=Δ

)1(11

ultnniit n

Kχψαρ ++

=

n

psn A

A=ρ

⎟⎟⎠

⎞⎜⎜⎝

⎛+=

n

pnnn I

eA 2

1α

ci

psi E

En =

21610*480 KKshsh γε −=

fhsstdsh kkkk=γ

tftk

citd +−

='461

Hkhs 0143.000.2 −=

( )735

941064 SV

ks

−=

cif f

k'1

5+

=

( )

⎟⎟⎠

⎞⎜⎜⎝

⎛++

⎟⎟⎠

⎞⎜⎜⎝

⎛−=

Δ+Δ−=

=Δ

124124

log55.045

31

i

t

py

popoi

po

pCRpSRi

itiipR

tt

fff

L

fff

KLf

φ

φ

cgpipES fnf =Δ

143

Figure A-6 Prestress Loss/Cracking Load Calculations according to the NCHRP 496 procedure for Specimen CC-70-2, page 2 of 3

Figure A-7 Prestress Loss/Cracking Load Calculations according to the NCHRP 496 procedure for Specimen CC-70-2, page 3 of 3

fcbi= 3.85 ksi

ΔP= 120.69 kipsΔfcb= 0.49 ksi

Mcr= 21288.11 in-kipsPcr= 137.9 kips

Cracking Load:Equations

Predicted Values( )

( )rcbcbibtt

ttcr

tt

bttptt

ttcb

pspESpT

ti

btig

ti

btipti

tiicbi

fffyI

M

IyPe

APf

AffP

IyM

Iye

APf

+Δ−=

Δ+Δ=Δ

Δ−Δ=Δ

−⎟⎟⎠

⎞⎜⎜⎝

⎛+=

*

1

144

Figure A-8 Prestress Loss/Cracking Load Calculations according to the AASHTO-LRFD Interim 2008 procedure for Specimen CC-70-2, page 1 of 3


Test: 5/27/08

INPUT

Gross Net Transform Gross Net Transform Number Height (in)A (in²) 494.9 492.78 528.77 494.9 491.5 516.4 8 at 2yb (in) 17.09 17.36 16.59 17.09 17.31 16.76 8 at 4I (in4) 82602 83403 87807 82602 82862 85936 8 at 6

As (in2) 0.612 0 0 0.612 0 0 6 at 8

ys (in) 38.75 38.75 38.75 38.75 38.75 38.75 2 at 10Aps_1 (in²/N) 0.153 0 0 0.153 0 0 2 at 12

N 36 36 36 36 36 36 2 at 14Aps (in

2) 5.508 0 0 5.508 0 0 C.G. 6.00ep (in) 11.09 11.36 10.59 11.09 11.31 10.76yp (in) 6.00 6.00 6.00 6.00 6.00 6.00

Ec (ksi) 4438 4438 4438 6387 6387 6387Eps (ksi) 29000 29000 29000 28809 28809 28809

n 6.53 6.53 6.53 4.51 4.51 4.51

1.0 fpi (ksi) 202.5 L (ft) 56.45833333Aps (in

2) 5.508 L span (ft) 555.36 fpy (ksi) 243 H (%) 7011.1 fpu (ksi) 270 V (in3) 335294.75

SA (in2) 85220.50.790 Pi (kips) 1115.4 V/SA (in) 3.9

04438 Δεp= 0.006986387 Δεp-Δεp_ES= 0.00626

150516

Mg_span (in-k) 2341.42467.2Mg_max (in-k)

fr_t (ksi)

Ect (ksi)Ec_t (ksi)

w (pcf)w (plf)

Measured fr_t (ksi)

Loss of Prestress/Cracking Load Calculations - Series 1 BeamMethod: AASHTO LRFD

Section PropertiesSection at Midspan

TermAt Release At Static Test Centerline

Strand Pattern

K1

f'ci (ksi)f'c_t (ksi)


ccc fwKE '000,33 5.11=

145


OUTPUT

0.9* Pi= 1003.83 kfcgp= 3.19 ksi fpy= 243.00

ΔfpES= 20.9 ksi fpt= 181.65KL= 30.00

Initial (fpi) 202.5 ksi ΔfpR= 1.2 ksiAfter ES (fpt) 181.65 ksi

Epst= 28809 ksi

esh= 0.000249628Kit= 0.79

ktd= 0.65khs= 1.02ks= 1.00kf= 0.79

ΔfpSR= 5.8 ksi

Quantitiy At Test At Ultimatektd= 0.65 1.00ks= 1.00 1.00 Kit= 0.79

khc= 1.00 1.00 Ψb(tt,ti)= 0.97kf= 0.79 0.79

Ψb(t,ti)= 0.97 1.49

ΔfpCR= 16.1 ksi



Equations Equations Stress

Total Losses:Elastic Shortening (ES)ShrinkageCreepRelaxation

Creep:

Relaxation:Equations Summary Equations Stress

Prestressing Force

Elastic Shortening:


cgpfctEpE

pESf =Δ

gIpgeg_M

gIpge

gAi*P.cgpfmax

21

90 −⎟⎟⎟⎟

⎠

⎞

⎜⎜⎜⎜

⎝

⎛

+=

itpshpSR KEf ε=Δ

310*48.0 −= tdfhsssh kkkkε

( )Hkhs 014.00.2 −=

( )SVk s 13.045.1 −=

cif f

k'1

5+

=

⎟⎟⎠

⎞⎜⎜⎝

⎛+−

=tf

tkci

td '461

[ ]),(7.0111

12

ifbg

pgg

g

ps

ci

pit

ttIeA

AA

EE

K

ψ+⎟⎟⎠

⎞⎜⎜⎝

⎛++

=

( ) 118.09.1, −= itdfhcsifb tkkkkttψ

Hkhc 008.056.1 −=

( )SVk s 13.045.1 −=

cif f

k'1

5+

=

⎟⎟⎠

⎞⎜⎜⎝

⎛+−

=tf

tkci

td '461

ititbcgpci

ppCR Kttf

EE

f ),(ψ=Δ

⎟⎟⎠

⎞⎜⎜⎝

⎛−=Δ 55.0

py

pt

L

ptpR f

fKf

f

146


feff= 158.64 ksiPeff= 873.80 kips

Mcr= 19702 in-kPcr= 131.3 kips


Predicted Values

⎟⎟⎠

⎞⎜⎜⎝

⎛+−+= r

g

bgtestg

g

bgcleff

g

eff

bg

gcr f

IyM

IyeP

AP

yI

M _

Figure A-11 Sample Box Beam Shop Drawing for specimen BB-01

147

Figure A-12 Sample Stress Calculations at Prestress Release for beam BB-01

Beam Mark: BB-01Series: Box

f'ci (psi) 4140 4B28 fcgp (ksi) 2.32Eci (ksi) 3307 Ag (in²) 678.8 n 8.77 Number Height (in) Number Height (in) Number Distance (ft) C.G. Eccentricty Po (kips)

K1 0.90 Ig (in4) 68745 ES (ksi) 20.32 22 at 2.5 22 at 2.5 4 10 3.27 10.35 725

K2 1.00 yb (in) 13.62 fpo (ksi) 182.18 6 at 4.5 6 at 4.5 4 4 3.41 10.21 613Eps (ksi) 29000 yt (in) 14.38 Po (kips) 836 2 at 6.5 2 at 6.5

fpi (ksi) 202.5 ecl (in) 10.45Strand Dia. (in) 0.5 eend (in) 10.45

Aps_1 (in²/N) 0.153 wu (k/ft) 0.707Ntotal 30 L (ft) 39

C.G. 3.17 C.G. 3.17Eccentricity 10.45 Eccentricity 10.45

x e Mg

ft ksi in ksi ksi in-k ksi ksiEnd 0.0 0.00 613 0.90 10.21 1.24 1.31 0 0.00 0.00Transfer Length 2.5 0.06 613 0.90 10.21 1.24 1.31 387.1 0.08 0.08

4 4.0 0.10 613 0.90 10.21 1.24 1.31 593.9 0.12 0.124.01 4.0 0.10 725 1.07 10.35 1.49 1.57 595.2 0.12 0.12

6 6.0 0.15 725 1.07 10.35 1.49 1.57 839.9 0.17 0.188 8.0 0.21 725 1.07 10.35 1.49 1.57 1052.0 0.21 0.22

10 10.0 0.26 725 1.07 10.35 1.49 1.57 1230.2 0.24 0.2610.01 10.0 0.26 836 1.23 10.45 1.73 1.83 1231.0 0.24 0.26

12 12.0 0.31 836 1.23 10.45 1.73 1.83 1374.4 0.27 0.2914 14.0 0.36 836 1.23 10.45 1.73 1.83 1484.7 0.29 0.31

14.5 14.5 0.37 836 1.23 10.45 1.73 1.83 1507.0 0.30 0.32Critical 17.0 0.44 836 1.23 10.45 1.73 1.83 1586.5 0.31 0.33Midspan 19.5 0.50 836 1.23 10.45 1.73 1.83 1613.0 0.32 0.34

ksi % of f'ci ksiEnd -2.14 51.8 0.41 6.31 OKTransfer Length -2.07 49.9 0.33 5.06 OK

4 -2.03 48.9 0.28 4.38 OK4.01 -2.44 58.8 0.38 5.86 OK

6 -2.39 57.7 0.33 5.06 OK8 -2.35 56.7 0.28 4.37 OK

10 -2.31 55.8 0.24 3.79 OK10.01 -2.72 65.7 0.34 5.27 OK

12 -2.69 65.0 0.31 4.80 OK14 -2.67 64.5 0.29 4.45 OK

14.5 -2.67 64.4 0.28 4.37 OKCritical -2.65 64.0 0.26 4.11 OKMidspan -2.64 63.9 0.26 4.03 OK


Section


LRFD Check

Section

Concrete

Centerline EndMaterial Properties Section Properties Elastic Shortening Strand Pattern

New ValueDebonding Pattern


Stress Calculations at Prestress Release - Box Beams

Steel

Lx

APo I

eyP boI

eyP toI

yM bgI

yM tg

cf '*

IyM

IeyP

AP

f bgboobot −+=

IyM

IeyP

AP

f tgtootop +−=

g

pg

g

p

gicgp I

eMI

eA

Pf −⎟⎟⎠

⎞⎜⎜⎝

⎛+=

219.0

cc

c ff

KKE '1000

'140.033000

5.1

21 ⎟⎠⎞

⎜⎝⎛ +=

oP

148

149

Figure A-13 Prestress Loss/Cracking Load Calculations according to the NCHRP 496 procedure for Specimen BB-01, page 1 of 3


Test: 9/18/08

INPUT

Gross Net Transform Gross Net Transform Number Height (in)A (in²) 678.8 688.66 728.91 678.8 681.5 704.1 22 at 2.5yb (in) 13.62 13.95 13.35 13.62 13.82 13.48 6 at 4.5I (in4) 68745 70236 74805 68745 69276 71812 2 at 6.5

As (in2) 1.86 0 0 1.86 0 0 at

ys (in) 26 26 26 26 26 26 atAps_1 (in²/N) 0.153 0 0 0.153 0 0 at

N 30 30 30 30 30 30 atAps (in

2) 4.59 0 0 4.59 0 0 C.G. 3.17ep (in) 10.45 10.78 10.19 10.45 10.66 10.31yp (in) 3.17 3.17 3.17 3.17 3.17 3.17

Ec (ksi) 3307 3307 3307 5876 5876 5876Eps (ksi) 29000 29000 29000 28884 28884 28884

n 8.77 8.77 8.77 4.92 4.92 4.92

K1 0.9 fpi (ksi) 202.5 Length (ft) 39K2 1.0 Aps (in2) 4.59 Span (ft) 36

f'ci (ksi) 4.14 fpy (ksi) 243 L (ft) 39.00f'c_t (ksi) 11.3 fpu (ksi) 270 H (%) 70

V (in3) 317678.4fr_t (ksi) 0.797 Pi (kips) 929.5 SA (in2) 72131.63

V/SA (in) 4.4Eci (ksi) 3307 Δεp= 0.00698276

Ec_t (ksi) 5876 Δεp-Δεp_ES= 0.00627365

w (pcf) 150w (plf) 707

Mg_span (in-k) 1374.408Mg_max (in-k) 1613.021

Loss of Prestress/Cracking Load Calculations - Box BeamMethod: NCHRP Report 496




28_'*5.7 cr ff =

cc

c ff

KKE '1000

'140.033000

5.1

21 ⎟⎠⎞

⎜⎝⎛ +=

150


OUTPUT

fcgp= 2.35 ksiΔfpES= 20.6 ksi Kit= 0.80

φi= 0.79Li= 1.15

Initial (fpi) 202.5 ksi ΔfpR= 0.7 ksiAfter ES (fpo) 181.94 ksi

Epst= 28884 ksi

esh= 0.000144K1= 0.9 Kit= 0.80K2= 1 ρn= 0.01

αn= 2.14ni= 8.77

γsh= 0.33 ψult= 1.47ktd= 0.39 χ= 0.7khs= 1.00ks= 0.88 ΔfpSR= 3.3 ksikf= 0.97

Quantitiy At Test At Ultimateγcr= 0.33 0.86ktd= 0.39 1.00 Kit= 0.80kla= 1.00 1.00 ρn= 0.01ks= 0.88 0.88 αn= 2.14

khc= 1.00 1.00 ni= 8.77kf= 0.97 0.97 ψult= 1.47

K1= 0.9 0.9 ψt= 0.57K2= 1 1 χ= 0.7ψt= 0.57 1.47

ΔfpCR= 9.3 ksi



Elastic Shortening: Relaxation:Equations Summary Equations Stress

Prestressing Force


Creep:Equations Equations Stress

Relaxation

Total Losses:Elastic Shortening (ES)ShrinkageCreep

cgpipES fnf =Δ

ci

psi

n

pnnn

n

psn

ultnniit

ittcgpipCR

EE

n

IeA

AA

nK

Kfnf

=

⎟⎟⎠

⎞⎜⎜⎝

⎛+=

=

++=

=Δ

2

1

)1(11

α

ρ

χψαρ

ψ

( )

⎟⎟⎠

⎞⎜⎜⎝

⎛++

⎟⎟⎠

⎞⎜⎜⎝

⎛−=

Δ+Δ−=

=Δ

124124

log55.045

31

i

t

py

popoi

po

pCRpSRi

itiipR

tt

fff

L

fff

KLf

φ

φ

itpsshpSR KEf ε=Δ

)1(11

ultnniit n

Kχψαρ ++

=

n

psn A

A=ρ

⎟⎟⎠

⎞⎜⎜⎝

⎛+=

n

pnnn I

eA 2

1α

ci

psi E

En =

21610*480 KKshsh γε −=

fhsstdsh kkkk=γ

tftk

citd +−

='461

Hkhs 0143.000.2 −=

( )735

941064 SV

ks

−=

cif f

k'1

5+

=

2190.1),( KKtt cri γψ =

fhcslatdcr kkkkk=γ

118.0−= ila tk

Hkhc 008.056.1 −=

tftk

citd +−

='461

( )735

941064 SV

ks

−=

cif f

k'1

5+

=

151


fcbi= 2.68 ksi

ΔP= 61.37 kipsΔfcb= 0.21 ksi

Mcr= 17412.62 in-kipsPcr= 187.2 kips

Cracking Load:

Predicted Values

Equations

( )

( )rcbcbibtt

ttcr

tt

bttptt

ttcb

pspESpT

ti

btig

ti

btipti

tiicbi

fffyI

M

IyPe

APf

AffP

IyM

Iye

APf

+Δ−=

Δ+Δ=Δ

Δ−Δ=Δ

−⎟⎟⎠

⎞⎜⎜⎝

⎛+=

*

1

152

Figure A-16 Prestress Loss/Cracking Load Calculations according to the AASHTO-LRFD Interim 2008 procedure for Specimen BB-01, page 1 of 3


Test: 9/18/08

INPUT

Gross Net Transform Gross Net Transform Number Height (in)A (in²) 678.8 687.71 725.63 678.8 681.6 704.4 22 at 2.5yb (in) 13.62 13.93 13.37 13.62 13.82 13.48 6 at 4.5I (in4) 68745 70113 74411 68745 69289 71851 2 at 6.5

As (in2) 1.86 0 0 1.86 0 0 at

ys (in) 26 26 26 26 26 26 atAps_1 (in²/N) 0.153 0 0 0.153 0 0 at

N 30 30 30 30 30 30 atAps (in

2) 4.59 0 0 4.59 0 0 C.G. 3.17ep (in) 10.45 10.77 10.20 10.45 10.66 10.31yp (in) 3.17 3.17 3.17 3.17 3.17 3.17

Ec (ksi) 3511 3511 3511 5800 5800 5800Eps (ksi) 29000 29000 29000 28802 28802 28802

n 8.26 8.26 8.26 4.97 4.97 4.97

0.9 fpi (ksi) 202.5 L (ft) 39Aps (in

2) 4.59 L span (ft) 364.14 fpy (ksi) 243 H (%) 7011.3 fpu (ksi) 270 V (in3) 317678.4

SA (in2) 72131.631160.797 Pi (kips) 929.5 V/SA (in) 4.4

3511 Δεp= 0.006985800 Δεp-Δεp_ES= 0.00632

150707

Mg_span (in-k) 1374.4081613.0

w (pcf)w (plf)

Mg_max (in-k)

Material PropertiesSteel AdditionalEquations Concrete

Ect (ksi)Ec_t (ksi)

K1

f'ci (ksi)f'c_t (ksi)


Loss of Prestress/Cracking Load Calculations - Box BeamMethod: AASHTO LRFD


fr_t (ksi)

ccc fwKE '000,33 5.11=

153


OUTPUT

0.9* Pi= 836.53 kfcgp= 2.32 ksi fpy= 243.00

ΔfpES= 19.1 ksi fpt= 183.36KL= 30.00

Initial (fpi) 202.5 ksi ΔfpR= 1.3 ksiAfter ES (fpt) 183.36 ksi

Epst= 28802 ksi

esh= 0.000184089Kit= 0.79

ktd= 0.39khs= 1.02ks= 1.00kf= 0.97

ΔfpSR= 4.2 ksi

Quantitiy At Test At Ultimatektd= 0.39 1.00ks= 1.00 1.00 Kit= 0.79

khc= 1.00 1.00 Ψb(tt,ti)= 0.71kf= 0.97 0.97

Ψb(t,ti)= 0.71 1.85

ΔfpCR= 10.8 ksi



Equations Stress

Creep:Equations Stress

Equations


Summary

Relaxation:

Equations

Elastic Shortening (ES)

Prestressing Force

Elastic Shortening:

Total Losses:

ShrinkageCreepRelaxation

cgpfctEpE

pESf =Δ

gIpgeg_M

gIpge

gAi*P.cgpfmax

21

90 −⎟⎟⎟⎟

⎠

⎞

⎜⎜⎜⎜

⎝

⎛

+=

itpshpSR KEf ε=Δ

310*48.0 −= tdfhsssh kkkkε

( )Hkhs 014.00.2 −=

( )SVk s 13.045.1 −=

cif f

k'1

5+

=

⎟⎟⎠

⎞⎜⎜⎝

⎛+−

=tf

tkci

td '461

[ ]),(7.0111

12

ifbg

pgg

g

ps

ci

pit

ttIeA

AA

EE

K

ψ+⎟⎟⎠

⎞⎜⎜⎝

⎛++

=

( ) 118.09.1, −= itdfhcsifb tkkkkttψ

Hkhc 008.056.1 −=

( )SVk s 13.045.1 −=

cif f

k'1

5+

=

⎟⎟⎠

⎞⎜⎜⎝

⎛+−

=tf

tkci

td '461

ititbcgpci

ppCR Kttf

EE

f ),(ψ=Δ

⎟⎟⎠

⎞⎜⎜⎝

⎛−=Δ 55.0

py

pt

L

ptpR f

fKf

f

154


feff= 167.10 ksiPeff= 766.99 kips

Mcr= 16370 in-kPcr= 176.0 kips

Predicted Values


⎟⎟⎠

⎞⎜⎜⎝

⎛+−+= r

g

bgtestg

g

bgcleff

g

eff

bg

gcr f

IyM

IyeP

AP

yI

M _

2)12(5

2

2

−=

LM

P crcr

155

APPENDIX B Concrete Batch Tickets from Type-C Beam and 4B28 Box

Beam Fabrication

Appendix B contains the information from the batch tickets used to fabricate all 55 girders in this research. The batch tickets provide the detailed proportions if all concrete mixture materials accounting for moisture content. The following list outlines the order of batch tickets presented in this appendix:

• Fabricator A (3 castings) • Fabricator B (2 castings) • Fabricator C (3 castings) • Fabricator D (4 castings) • Fabricator E (4 castings)

156

Fabricator A Batch Tickets:

Material Sources

Fine Aggregate TXI-Green Pit Cement Alamo Type III

Coarse Aggregate Hanson-Ogden Retarding Admixture Sika Plastiment

Water San Marcos Water Co High Range Water Reducer Sika Viscocrete 2100

Fabricator A: Casting #1

Date: 9/26/07 w/cm ratio: 0.368

Batch Size: 4 yards Mix Design: 21P- ATH6.0

Material Description Actual Mix Target Value

Fine Aggregate 6066 lbs 6107 lbs

Coarse Aggregate 7213 lbs 7213 lbs

Water 656 lbs 660 lbs

Cement 2240 lbs 2256 lbs

Retarding Admixture 68 oz 68 oz

High Range Water Reducer 161 oz 160 oz


Date: 9/26/07 w/cm ratio: 0.373










Date: 9/26/07 w/cm ratio: 0.369










Date: 9/26/07 w/cm ratio: 0.373










Date: 9/26/07 w/cm ratio: 0.369










Date: 9/26/07 w/cm ratio: 0.365









157


Date: 10/03/07 w/cm ratio: 0.355










Date: 10/03/07 w/cm ratio: 0.357










Date: 10/03/07 w/cm ratio: 0.356










Date: 10/03/07 w/cm ratio: 0.352










Date: 10/03/07 w/cm ratio: 0.358










Date: 10/03/07 w/cm ratio: 0.354










Date: 10/03/07 w/cm ratio: 0.357










Date: 10/03/07 w/cm ratio: 0.351









158


Date: 10/03/07 w/cm ratio: 0.358










Date: 10/03/07 w/cm ratio: 0.351










Date: 10/09/07 w/cm ratio: 0.357










Date: 10/09/07 w/cm ratio: 0.363










Date: 10/03/07 w/cm ratio: 0.354










Date: 10/03/07 w/cm ratio: 0.362










Date: 10/09/07 w/cm ratio: 0.357










Date: 10/09/07 w/cm ratio: 0.361









159


Date: 10/09/07 w/cm ratio: 0.363










Date: 10/09/07 w/cm ratio: 0.369










Date: 10/09/07 w/cm ratio: 0.362










Date: 10/09/07 w/cm ratio: 0.365










Date: 10/09/07 w/cm ratio: 0.364










Date: 10/09/07 w/cm ratio: 0.369










Date: 10/09/07 w/cm ratio: 0.370










Date: 10/09/07 w/cm ratio: 0.355









160

Fabricator B Batch Tickets:

Fabricator B: Casting #1

Date: 4/23/08 w/cm ratio: 0.27

Batch Size: 3.5 yards Mix Design: 437-2.08






Class F Fly Ash 575 lbs N/A lbs











Class F Fly Ash 580 lbs 578 lbs

























Material Sources

Fine Aggregate Arena Cement Alamo Type III 1a

Coarse Aggregate Eagle Lake Retarding Admixture Plastiment

Water Well Water High Range Water Reducer Sika 2100

Class F Fly Ash Headwaters Jewitt, TX

161



































































162



































































163



































































164



































































165

Fabricator C Batch Tickets:

Material Sources

Fine Aggregate River Sand Multisources Cement Capital Type III

Coarse Aggregate Vulcan 1604 Grade 5 Retarding Admixture Sika Plastiment

Water BCW Well High Range Water Reducer Sika 2100

Fabricator C: Casting #1


Batch Size: 4 yards Mix Design: 7S2100C




Water 618.0 lbs 616.4 lbs






















































166

















































































167

















Water 2656 lbs 641.8 lbs

Cement 642.0 lbs 2632 lbs































































168































169

Fabricator D Batch Tickets:

Material Sources

Fine Aggregate Fordyce Murphy Cement Alamo Type III

Coarse Aggregate Fordyce Murphy Retarding Admixture Pozzolith 300R

Water City of Victoria High Range Water Reducer Rheobuild 1000

Fabricator D: Casting #1


Batch Size: 3.75 yards Mix Design: PT-SWR-SS






Retarding Admixture 69.0 oz 68.6 oz











High Range Water Reducer oz 619 oz









































170







































































**Batch tickets for castings 3 and 4 were not

available**

171

Fabricator E Batch Tickets:

Castings 1,3,and 4 Material Sources

Manufactured Fine Aggregate Hanson Type III Cement Alamo Type III

Natural Fine Aggregate TXI Green Pit Class F Fly Ash Headwaters Jewitt, TX

Limestone Coarse Aggregate Hanson-Ogden Quarry Retarding Admixture Sika Plastiment

River Gravel Coarse Aggregate TXI River Gravel High Range Water Reducer Sika Viscocrete 4100

Water San Marcos Water Co.

Casting 2 Material Sources

Fine Aggregate Wrights FM 3088 Class F Fly Ash Headwaters Jewitt, TX

Coarse Aggregate Wrights Realitos Retarding Admixture Sika Plastiment

Type I/II Cement Alamo Type I/II LA High Range Water Reducer Sika Viscocrete 2100

Water City of Corpus Christi

Casting Concrete Type Coarse Aggregate Type Date

1 Conventional Limestone 8/21/08

2 Conventional River Gravel 7/22/08

3 SCC Limestone 8/25/08

4 SCC River Gravel 8/26/08

Fabricator E: Casting #1

Date: 8/21/08 w/cm ratio: 0.295

Batch Size: 2.5 yards Mix Design:8.5-25-4100










Date: 8/21/08 w/cm ratio: 0.292

Batch Size: 3 yards Mix Design:8.5-25-4100










Date: 8/21/08 w/cm ratio: 0.295











Date: 8/21/08 w/cm ratio: 0.294










172


Date: 8/21/08 w/cm ratio: 0.295











Date: 7/22/08 w/cm ratio: 0.273

Batch Size: 2.8 yards Mix Design: 22.06










Date: 7/22/08 w/cm ratio: 0.273

Batch Size: 3 yards Mix Design: 22.06










Date: 8/21/08 w/cm ratio: 0.292

Batch Size: 2.5 yards Mix Design:8.5-25-4100










Date: 7/22/08 w/cm ratio: 0.273











Date: 7/22/08 w/cm ratio: 0.273










173


Date: 7/22/08 w/cm ratio: 0.273











Date: 7/22/08 w/cm ratio: 0.273











Date: 8/25/08 w/cm ratio: 0.354

Batch Size: 2.5 yards Mix Design: SCC7.5UT










Date: 7/22/08 w/cm ratio: 0.273











Date: 7/22/08 w/cm ratio: 0.273











Date: 8/25/08 w/cm ratio: 0.356

Batch Size: 3 yards Mix Design: SCC7.5UT









174


Date: 8/25/08 w/cm ratio: 0.355











Date: 8/25/08 w/cm ratio: 0.336











Date: 8/26/08 w/cm ratio: 0.354

Batch Size: 2.5 yards Mix Design: SCC7.45S










Date: 8/25/08 w/cm ratio: 0.346











Date: 8/25/08 w/cm ratio: 0.347

Batch Size: 1.5 yards Mix Design: SCC7.5UT










Date: 8/26/08 w/cm ratio: 0.357










175


Date: 8/26/08 w/cm ratio: 0.362

Batch Size: 3 yards Mix Design: SCC7.45S










Date: 8/26/08 w/cm ratio: 0.354











Date: 8/26/08 w/cm ratio: 0.359











Date: 8/26/08 w/cm ratio: 0.356











Date: 8/26/08 w/cm ratio: 0.357










176

177

APPENDIX C Static Flexural Test Specimen Information

Appendix C contains the detailed testing information for each of the 55 girders tested in

flexure. The table in this appendix presents the specimen identification information, fabrication and test dates, maximum compressive stresses at prestress transfer, measured cracking loads, predicted cracking loads, and predicted cracking load accuracies using both the NCHRP 496 prestress loss procedure and the AASHTO-LRFD (2008) prestress loss procedure.

Figure C-1 Static Flexural Test Information

Concrete Strength at Test

Test Mark PrecasterPrecaster

Mark SeriesDesign Release

StrengthCoarse Aggregate Date of Cast Date of Release Date of Test

Actual Release Strength, psi Hold Down

Critical Section Transfer Length Strength, psi

NCHRP Method

NCHRP with Dead Loads

AASHTO-LRFD Method

AASHTO-LRFD with DL

Measured Pcr, kips

NCHRP Error (%)

AASHTO-LRFD Error (%)

1 CA-70-3 A A1 1 4000 Limestone 9/26/2007 9/26/2007 11/1/2007 3940 0.703 0.700 0.702 10200 109 108.228 103 102.228 107 -1.15 4.462 CA-70-2 A A2 1 4000 Limestone 9/26/2007 9/26/2007 11/15/2007 3930 0.705 0.702 0.703 10550 109 108.228 103 102.228 110 1.61 7.073 CA-70-1 A A3 1 4000 Limestone 9/26/2007 9/26/2007 11/27/2007 3930 0.705 0.702 0.704 10800 108 107.228 102 101.228 108 0.71 6.274 CA-60-1 A A6 1 4700 Limestone 10/3/2007 10/4/2007 11/29/2007 4540 0.616 0.613 0.614 10500 110 109.228 104 103.228 111 1.60 7.005 CA-60-2 A A4 1 4700 Limestone 10/3/2007 10/4/2007 12/4/2007 4540 0.616 0.613 0.614 10700 110 109.228 104 103.228 112 2.48 7.836 CA-60-3 A A5 1 4700 Limestone 10/3/2007 10/4/2007 12/12/2007 4540 0.616 0.613 0.614 11050 110 109.228 103 102.228 111 1.60 7.907 CA-65-1 A A3 1 4300 Limestone 10/9/2007 10/10/2007 12/18/2007 4370 0.638 0.636 0.636 10200 108 107.228 102 101.228 110 2.52 7.978 CA-65-2 A A6 1 4300 Limestone 10/9/2007 10/10/2007 1/4/2008 4380 0.637 0.634 0.635 11150 108 107.228 103 102.228 107 -0.21 4.469 CA-65-3 A A4 1 4300 Limestone 10/9/2007 10/10/2007 1/8/2008 4310 0.647 0.644 0.645 11400 108 107.228 102 101.228 113 5.11 10.42

10 CA-65-4 A A5 1 4300 Limestone 10/9/2007 10/10/2007 1/10/2008 4340 0.641 0.640 0.641 11500 108 107.228 102 101.228 110 2.52 7.9711 CA-65-5 A A2 1 4300 Limestone 10/9/2007 10/10/2007 1/15/2008 4330 0.644 0.641 0.642 11800 108 107.228 102 101.228 111 3.40 8.8012 CA-65-6 A A1 1 4300 Limestone 10/9/2007 10/10/2007 1/17/2008 4300 0.648 0.645 0.646 11900 108 107.228 102 101.228 109 1.63 7.1313 CD-70-1 D D1 2 5400 River Gravel 3/4/2008 3/5/2008 4/2/2008 5580 0.683 0.681 0.679 11000 144 143.228 138 137.228 142 -0.86 3.3614 CD-70-2 D D2 2 5400 River Gravel 3/4/2008 3/5/2008 4/7/2008 5500 0.692 0.690 0.688 11600 143 142.228 137 136.228 143 0.54 4.7415 CD-70-3 D D3 2 5400 River Gravel 3/4/2008 3/5/2008 4/10/2008 5420 0.701 0.699 0.698 12000 142 141.228 138 137.228 147 3.93 6.6516 CD-65-1 D D1 2 5850 River Gravel 3/7/2008 3/8/2008 4/14/2008 5670 0.673 0.671 0.669 9600 142 141.228 135 134.228 143 1.24 6.1317 CD-65-2 D D2 2 5850 River Gravel 3/7/2008 3/8/2008 4/18/2008 5670 0.673 0.671 0.669 9600 141 140.228 134 133.228 135 -3.87 1.3118 CD-65-3 D D3 2 5850 River Gravel 3/7/2008 3/8/2008 4/21/2008 5670 0.673 0.671 0.669 9600 141 140.228 134 133.228 137 -2.36 2.7519 CD-65-4 D D1 2 5850 River Gravel 3/14/2008 3/15/2008 4/24/2008 5940 0.644 0.642 0.640 10700 143 142.228 136 135.228 136 -4.58 0.5720 CD-65-5 D D2 2 5850 River Gravel 3/14/2008 3/15/2008 4/29/2008 5940 0.644 0.642 0.640 11200 142 141.228 136 135.228 148 4.58 8.6321 CD-65-6 D D3 2 5850 River Gravel 3/14/2008 3/15/2008 5/1/2008 5940 0.644 0.642 0.640 11400 142 141.228 136 135.228 143 1.24 5.4322 CD-60-1 D D1 2 6400 River Gravel 3/12/2008 3/13/2008 5/5/2008 6320 0.608 0.606 0.604 11700 143 142.228 137 136.228 143 0.54 4.7423 CD-60-2 D D2 2 6400 River Gravel 3/12/2008 3/13/2008 5/12/2008 6310 0.609 0.607 0.605 12000 142 141.228 137 136.228 146 3.27 6.6924 CD-60-3 D D3 2 6400 River Gravel 3/12/2008 3/13/2008 5/20/2008 6300 0.610 0.608 0.606 12400 142 141.228 136 135.228 142 0.54 4.7725 CC-70-1 C A4 2 5400 Limestone 3/14/2008 3/15/2008 5/23/2008 5380 0.706 0.704 0.702 10700 138 137.228 131 130.228 142 3.36 8.2926 CC-70-2 C A5 2 5400 Limestone 3/14/2008 3/15/2008 5/27/2008 5360 0.709 0.706 0.705 11100 138 137.228 131 130.228 138 0.56 5.6327 CC-70-3 C A6 2 5400 Limestone 3/14/2008 3/15/2008 5/29/2008 5330 0.712 0.710 0.709 11300 138 137.228 131 130.228 139 1.27 6.3128 CC-65-1 C A6 2 5850 Limestone 3/26/2008 3/27/2008 5/29/2008 5970 0.641 0.639 0.637 11050 141 140.228 134 133.228 136 -3.11 2.0429 CC-65-2 C A3 2 5850 Limestone 3/26/2008 3/27/2008 6/2/2008 6000 0.638 0.636 0.634 11180 141 140.228 134 133.228 141 0.55 5.5130 CC-65-3 C A2 2 5850 Limestone 3/26/2008 3/27/2008 6/3/2008 6130 0.626 0.624 0.622 11200 141 140.228 135 134.228 135 -3.87 0.5731 CC-65-4 C A5 2 5850 Limestone 3/26/2008 3/27/2008 6/10/2008 6070 0.631 0.630 0.628 11450 140 139.228 134 133.228 134 -3.90 0.5832 CC-65-5 C A4 2 5850 Limestone 3/26/2008 3/27/2008 6/11/2008 6350 0.606 0.604 0.602 11480 141 140.228 135 134.228 141 0.55 4.8033 CC-65-6 C A1 2 5850 Limestone 3/26/2008 3/27/2008 6/12/2008 6250 0.615 0.613 0.611 11500 141 140.228 135 134.228 142 1.25 5.4734 CC-60-1 C A6 2 6400 Limestone 4/1/2008 4/2/2008 6/12/2008 6350 0.606 0.604 0.602 10750 141 140.228 135 134.228 130 -7.87 -3.2535 CC-60-2 C A4 2 6400 Limestone 4/1/2008 4/2/2008 6/17/2008 6370 0.604 0.602 0.600 10800 141 140.228 134 133.228 130 -7.87 -2.4836 CC-60-3 C A5 2 6400 Limestone 4/1/2008 4/2/2008 6/18/2008 6370 0.604 0.602 0.600 10800 141 140.228 134 133.228 138 -1.61 3.4637 CB-60-1 B L5 1 5850 River Gravel 4/23/2008 4/23/2008 6/19/2008 4820 0.588 0.585 0.585 12300 112 111.228 108 107.228 130 14.44 17.5238 CB-60-2 B L4 1 5850 River Gravel 4/23/2008 4/23/2008 6/23/2008 5010 0.567 0.564 0.564 12700 113 112.228 108 107.228 126 10.93 14.9039 CB-60-3 B L6 1 5850 River Gravel 4/23/2008 4/23/2008 6/24/2008 4620 0.612 0.609 0.609 12800 111 110.228 107 106.228 120 8.14 11.4840 CB-70-1 B G1 1 5400 River Gravel 5/6/2008 5/6/2008 6/25/2008 4540 0.622 0.619 0.619 12100 112 111.228 107 106.228 118 5.74 9.9841 CB-70-2 B G2 1 5400 River Gravel 5/6/2008 5/6/2008 6/26/2008 4360 0.646 0.644 0.644 12140 111 110.228 107 106.228 121 8.90 12.2142 CB-70-3 B G3 1 5400 River Gravel 5/6/2008 5/6/2008 6/27/2008 4180 0.673 0.670 0.670 12200 110 109.228 106 105.228 118 7.43 10.8243 CB-70-4 B G4 1 5400 River Gravel 5/6/2008 5/6/2008 7/1/2008 4030 0.696 0.693 0.694 12430 110 109.228 106 105.228 110 0.70 4.3444 CB-70-5 B G5 1 5400 River Gravel 5/6/2008 5/6/2008 7/2/2008 3880 0.721 0.718 0.719 12500 109 108.228 105 104.228 118 8.28 11.6745 CB-70-6 B G6 1 5400 River Gravel 5/6/2008 5/6/2008 7/7/2008 3680 0.758 0.755 0.756 12800 108 107.228 104 103.228 112 4.26 7.8346 BB-01 E D2 Box 4100 Limestone 8/21/2008 8/21/2008 9/18/2008 4140 0.640 11300 187 185.028 176 174.028 176 -5.13 1.1247 BB-02 E D1 Box 4100 Limestone 8/21/2008 8/21/2008 9/18/2008 4070 0.650 11300 187 185.028 176 174.028 178 -3.95 2.2348 BB-03 E M1 Box 4100 River Gravel 7/22/2008 7/22/2008 9/4/2008 4120 0.657 11160 183 181.028 176 174.028 195 7.17 10.7549 BB-04 E M3 Box 4100 River Gravel 7/22/2008 7/22/2008 8/19/2008 4220 0.643 10672 186 184.028 179 177.028 196 6.11 9.6850 BB-05 E M2 Box 4100 River Gravel 7/22/2008 7/22/2008 8/28/2008 4170 0.650 10900 184 182.028 178 176.028 200 8.99 11.9951 BB-06 E D2 Box 4100 Limestone 8/25/2008 8/26/2008 10/2/2008 4060 0.641 9530 183 181.028 167 165.028 162 -11.75 -1.8752 BB-07 E D1 Box 4100 Limestone 8/25/2008 8/26/2008 10/7/2008 4060 0.641 9575 182 180.028 166 164.028 155 -16.15 -5.8253 BB-08 E D3 Box 4100 River Gravel 8/26/2008 8/27/2008 9/24/2008 4040 0.655 8730 183 181.028 170 168.028 161 -12.44 -4.3754 BB-09 E D2 Box 4100 River Gravel 8/26/2008 8/27/2008 9/25/2008 4040 0.655 8900 183 181.028 170 168.028 162 -11.75 -3.7255 BB-10 E D1 Box 4100 River Gravel 8/26/2008 8/27/2008 9/30/2008 4040 0.655 9730 183 181.028 171 169.028 163 -11.06 -3.70

Specimen Information

Cracking LoadBottom Fiber Compressive Stress at Release

(*f'ci) Predicted Cracking Load

178

179

APPENDIX D Bursting Cracks in 4B28 Box Beam Specimens

Cracks forming due to bursting, spalling, and splitting stresses in the end regions

of the 4B28 box beams were monitored during this phase of TxDOT Project 5197. The observed cracks were termed “bursting cracks.” All cracks were marked on the beam and crack widths were measured using a crack comparator card. The observed crack widths in the box beams did not exceed 0.007-inches. Anything smaller than 0.005-inches was termed hairline. Both sides of each end of the beam were inspected and photographed totaling 40 photographs in all (10 specimens x four sides from each specimen). The crack maps for each end of the beam are presented in this appendix.

After inspecting all 10 box beams for bursting cracks, several observations were made. First, the most common bursting crack occurred at the re-entrant corner of the blockout at each end of the box beam. This bursting crack was the largest, widest, and most consistent crack among all 10 specimens. Secondly, the coarse aggregate type (limestone or hard river gravel) used in specimen fabrication did not appear to affect the number, length, or width of the bursting cracks. Finally, the use of conventional concrete or SCC mixture designs did not appear to effect the number, length or width of the bursting cracks.

The following pages present the crack maps for the bursting cracks in the box beams. All observed cracks wider than “hairline” are noted on the crack maps. For convenience, Figure D-1 provides a lettering key for the crack maps that follow.

Figure D-1 Bursting crack map key for box beams

180

Specimen BB-01

Specimen BB-02

Specimen BB-03

28”

1’-0”

A B DC

South End North End

0.005” 0.005”

28”

1’-0”

A B DC

South End North End

0.005” 0.005”

28”

1’-0”

A B DC

South End North End

181

Specimen BB-04

Specimen BB-05

Specimen BB-06

28”

1’-0”

A B DC

South End North End

28”

1’-0”

A B DC

South End North End

28”

1’-0”

A B DC

South End North End

0.005”0.005”

182

Specimen BB-07

Specimen BB-08

Specimen BB-09

28”

1’-0”

A B DC

South End North End

0.005”0.007” 0.005”

28”

1’-0”

A B DC

South End North End

0.005”0.005” 0.005”

28”

1’-0”

A B DC

South End North End

0.005”

183

Specimen BB-10

28”

1’-0”

A B DC

South End North End

0.005”

184

185

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